A few comments on pile design

Yves Robert

Abstract: Sixty-three static pile load tests were analyzed with respect to published soil properties to validate or refinewell-known pile design methods. Even though most of the data used to carry out this study are for driven piles, someconsideration was given to bored piles. Some comments were made concerning pile design in general.

Key words: bearing capacity, skin friction, end bearing, pile design, pile load test.

Rsum : Soixante trois essais de chargement statique de pieux ont t analyss en fonction des proprits des sols qui ont tpublies afin de valider ou de raffiner des mthodes de conception de pieux bien connues. Mme si la plupart des donnesutilises pour raliser la prsente tude correspondent des pieux battus, les pieux fors ont aussi t considrs. Descommentaires ont aussi t donns pour la conception des pieux en gnral.

Mots cls : capacit portante, rsistance en frottement, rsistance en pointe, conception des pieux, essai de chargement sur pieux.

Introduction

The ultimate bearing capacity of a pile (Qu) can be defined asthe sum of the skin resistance (QS) and the toe resistance (QP)as shown by

[1] Qu = QS + QP = fSi ASi + qP APwhere fSi is the unit skin friction on shaft segment i, ASi is thearea of this pile segment, qP is the unit toe resistance, and AP isthe toe area.

The unit toe resistance is often calculated using bearingcapacity formulas similar to the following:[2] qP = cNC + tNq + 0.5DNwhere c is the effective cohesion; D is the pile diameter; t isthe vertical effective stress at the pile toe; is the unit weightof the soil; and N, Nq, and Nc are bearing capacity factors. TheN term is usually neglected, and eqs. [3] and [4] are used tocompute the unit toe resistance in granular and cohesive soils,respectively:[3] qP = tNq[4] qP = cNC

The value of Nq is a function of the internal friction angle of the soil and of the assumed shape of the shear pattern nearthe pile toe. As shown by Coyle and Castello (1981) andCoduto (1994), there are major variations in the value of Nqpredicted by the different theories and choosing the right valueis usually difficult. The value of NC is generally close to 9.

The unit skin friction fSi can be calculated as[5] fSi = KiVi tan iwhere Ki is the coefficient of horizontal earth pressure, Vi is

the effective vertical stress at the centre of pile segment i, andi is the effective friction angle between the soil and the pilematerial at the location of this pile segment.

The engineer who uses eqs. [2][5] to design piles mustknow the total unit weight , the internal friction angle , theeffective cohesion c, the coefficient of horizontal earth pres-sure K, and the effective friction angle between the soil andthe pile material for each soil layers. These parameters cannotbe measured directly during a standard geotechnical investiga-tion in granular soils. They can, however, be estimated throughcorrelations with in situ test results and laboratory tests carriedout on reconstituted granular samples, since getting undis-turbed granular samples is rather difficult. Recovering intactclay samples is generally fairly easy, but the cost of measuringsome of those soil properties in the laboratory can often bedifficult to justify for small- and even medium-size projects.It is for these reasons that empirical design methods are oftenused to design piles.

Meyerhof (1976) linked both fs and qP (in kPa) to the stand-ard penetration test index N using eqs. [6] and [7]:

[6a] qP = 40NPTLD 400N (driven piles)

[6b] qP = 12NPTLD 120N (bored piles)

[7a] fs = 2NA (driven piles)[7b] fs = NA (low-displacement piles)where L is the penetration of the pile into a soil layer; NA is theaverage standard penetration resistance within the embeddedlength of the pile in uniform soil deposits or within each soillayer in stratified deposits; and NPT is the standard penetrationresistance, near the pile toe, corrected for an overburden pres-sure of 100 kPa using the method proposed by Peck et al.(1974). Meyerhof also used the concept of critical depth belowwhich the unit skin and toe resistances remain constant. Thiscritical-depth concept has also been used by Tavenas (1971)and Vesic (1970).

According to Fellenius (1994), the critical-depth concept isthe result of the neglect of residual loads in the test piles. Coyle

Received November 21, 1996. Accepted March 26, 1997.

Y. Robert. Quformat Ltd., 591 Le Breton, Longueuil, QCJ4G 1R9, Canada.

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and Castello (1981) found that, rather than becoming constantbelow a certain critical depth, the unit skin and toe resistanceswill continue to increase with increasing depth, although at adecreasing rate. Kraft (1991) discussed extensively the criti-cal-depth concept and concluded that it does not exist. TheCanadian foundation engineering manual (CanadianGeotechnical Society 1992) used to recommend application ofthe concept of critical depth when designing pile foundations.The third edition of the manual mentions, however, that theevidence available is sufficient to cast some doubt on the rele-vance of the critical-depth concept and caution should be usedin designing long piles in cohesionless soils.

It should be mentioned that residual loads are loads whichare present in the piles before any measurements are takenduring the load tests and neglecting them will cause an over-estimation of the skin resistance and an underestimation of thetoe resistance during a compressive load test. Furthermore,according to Kraft (1991), neglecting the residual loads haslittle effect on the computed axial capacity of piles 1525 mlong and might underpredict it in the cases of piles 7590 mlong.

In cohesive soils, the method is often used to design piles.It is a simple design method which uses the undrained shearstrength of the cohesive soil (Su) to calculate the unit skinresistance as shown by[8] fs = Suwhere is equal to 1 for low values of Su and decreases usinga nonlinear function to about 0.3 for high values of Su. Thismethod is very popular, since values of Su can be easily ob-tained from vane shear tests or unconfined compression tests.Equation [4] is normally used to compute the unit toe resis-tance in cohesive soil, with NC equal to 9 and c replaced by Su.

It should be pointed out that other pile design methods areavailable. They are often based on empirical relationships anduse data from laboratory tests or in situ testing devices such asthe static cone penetrometer, the piezocone, or the pressure-meter. It is, however, beyond the scope of this paper to reviewthese design methods.

Database

A database of 63 compression pile load tests was used to carryout this study. Fifty-three of these load tests were carried outon driven piles, and bored piles were used for the remainingload tests. The data for these load tests were gathered from theavailable literature. To be included in the database, a load testhad to have been carried out to failure on a pile with a uniformcross section. Some timber piles with their conical shape andcast in place concrete piles with their more or less uniformcross sections were also included because they are widelyused. The geotechnical data provided in each paper had to besufficient to allow the author to predict the bearing capacity ofthe test pile(s) using conventional design methods. However,some engineering judgement was usually required to completethe data provided in the papers.

To correlate soil and load test data for different test sites, asingle failure criterion must be used. The author chose to usethe Davisson (1972) failure criterion, since it is well knownand widely used. With this failure criterion, the ultimate bearingcapacity is defined as the load corresponding to the intersection

between the elastic deformation curve of the pile, shifted alongthe deformation axis by a value equal to 3.81 plus the diameter ofthe pile (in millimetres) divided by 120, and the load-deformationcurve recorded during the load test.

The Davisson capacity was computed by the author when-ever it was not provided and all load-deformation data wereincluded in the paper. In some cases, however, the Davissoncapacity had to be estimated using some load-deformation dataprovided in the paper or through a rough correlation betweenthe Davisson criterion and some other failure criterion. Theresults of these estimations were always within 10% of thepublished capacities. In the case of brittle failure, when theload reaches a maximum value after a small deformation anddecreases afterward, the maximum load applied to the pile wasused as the ultimate capacity. Table 1 shows a summary of theload tests used to build the database.

Pile design

In this study, the author did not try to match the published unitskin and toe resistances because the residual stresses locked inthe test piles probably hid their true distribution. Some of thepapers used to build the database are more than 20 years oldand the published resistance distributions may not be accurate.However, as mentioned in the Introduction, residual stresseshave little influence on the measured total capacity of pileswith length smaller than 75 m. It was therefore decided to onlytry to predict the total bearing capacity of the test piles andoptimize the different design equations to minimize the differ-ence between the computed capacity and the values given inTable 1.

The first step in predicting the bearing capacity of a pile isto divide the soil deposits into layers of constant geotechnicalproperties. The thickness of these soil layers will be a functionof the natural stratigraphy of the site and of the uniformity ofeach soil stratum. Each layer will have a constant unit weight,standard penetration resistance, or undrained shear strengthand will provide a constant unit skin friction.

The assumed shape of the shear pattern and its extent aboveand below the pile toe vary depending on the theory. It is fairlycertain, however, that the unit resistance in this part of the pilewill be a function of the geotechnical properties of the soilsfound above and below the pile toe. The author tried to predictthe bearing capacity of the test piles using the average toebearing capacity calculated with the geotechnical properties ofthe soils found 1.5 pile diameters above and below the pile toe,3 pile diameters above and below the pile toe, and within 4pile diameters below and 6 pile diameters above the pile toe.A shear area extending 3 pile diameters above and below thepile toe provided the best overall prediction. The average toebearing capacity was calculated by subdividing the soil depos-its near the pile toe into quarter-diameter-thick layers. Thegeotechnical properties of these soil layers were used to com-pute toe bearing capacity values, which were then averaged tocompute the actual toe bearing capacity of the test piles.

In some cases, however, when the pile toe comes close toa weaker soil stratum, a punching failure might occur. In thesecases, whenever the unit end bearing capacity, within 3 pilediameters below the pile toe, drops to a value that is less than55% of the value found at the pile toe, the lower bearing valuewill be used until the shear strength of the soils starts to increase

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Authors Pile type and dimensionsEmbeddeddepth (m)

Soil type alongthe shaft

Soil typeat the toe

Ultimatecapacity (kN)

Ahmed and Sowers 1985 457 mm square concrete 41.30 Cohesive Cohesive 1872Blanchet et al. 1980 Closed-bottom, 219 mm pipe 23.80 Cohesive Cohesive 390Blanchet et al. 1980 Herkules H420 23.80 Cohesive Cohesive 585Blanchet et al. 1980 Hercules H420 37.50 Cohesive Cohesive 845Blanchet et al. 1980 368 mm timber 15.25 Cohesive Cohesive 698Blanchet et al. 1980 374 mm timber 15.25 Cohesive Cohesive 640Bowles 1977 Closed-bottom, 356 mm pipe 15.24 Granular Granular 1212Canadian Wood Council 1991 360 mm timber 12.70 Granular Granular 670Endley et al. 1979 Closed-bottom, 273 mm pipe 32.00 Cohesive Cohesive 1560FHWA 1991 457 mm square concrete 19.75 Granular Granular 1667FHWA 1991 457 mm square concrete 23.00 Granular Granular 2540FHWA 1991 Hollow, 610 mm square concrete 19.66 Granular Granular 2870FHWA 1991 Hollow, 610 mm square concrete 22.92 Granular Granular 3726FHWA 1991 Hollow, 914 mm square concrete 22.25 Granular Granular 4903Ismael and Klym 1979 1070 mm bored 6.40 Granular Granular 1783Jaime and Romo 1990 310 mm square concrete 15.00 Cohesive Cohesive 540Laier 1994 12-H74 36.30 Granular Granular 2000Mansur and Hunter 1970 406 mm square concrete 12.25 Granular Granular 1667Mansur and Hunter 1970 14-BP73 16.21 Granular Granular 1917Mansur and Hunter 1970 14-BP73 12.19 Granular Granular 1310aMansur and Hunter 1970 380 mm timber 11.77 Granular Granular 731Mansur and Hunter 1970 14-BP73 15.88 Granular Granular 1646McCammon and Golder 1970 610 mm, open-bottom pipe 45.40 Granular Granular 1560McCammon and Golder 1970 610 mm, closed-bottom pipe 47.24 Granular Granular 3967McCammon and Golder 1970 610 mm, open-bottom pipe 30.48 Cohesive Cohesive 1961McCammon and Golder 1970 610 mm, closed-bottom pipe 48.15 Cohesive Cohesive 3495Nordlund 1963 324 mm, closed-bottom pipe 24.40 Granular Granular 1114Nordlund 1963 324 mm, closed-bottom pipe 16.80 Granular Granular 445Nordlund 1963 324 mm, closed-bottom pipe 18.30 Granular Granular 845ONeill and Reese 1972 763 mm bored 7.00 Cohesive Cohesive 1150ONeill and Reese 1972 763 mm bored 14.00 Cohesive Cohesive 2550ONeill and Reese 1972 763 mm bored 7.00 Cohesive Cohesive 1070Parsons 1966 203 mm, closed-bottom pipe 14.30 Granular Granular 392Reese et al. 1976 850 mm bored 18.30 Cohesive Cohesive 6660aReese et al. 1976 762 mm bored 7.74 Granular Granular 4010aReese et al. 1976 762 mm bored 22.40 Cohesive Granular 5535aTavenas 1971 Herkules H800 21.00 Granular Granular 1235Tavenas 1971 Herkules H800 18.00 Granular Granular 1040Tavenas 1971 Herkules H800 14.94 Granular Granular 929Tavenas 1971 Herkules H800 11.89 Granular Granular 691Tavenas 1971 Herkules H800 8.81 Granular Granular 490Tavenas 1971 Herkules H800 5.80 Granular Granular 312Tavenas 1971 12-BP74 20.75 Granular Granular 1872Tavenas 1971 12-BP74 17.70 Granular Granular 1003Tavenas 1971 12-BP74 11.60 Granular Granular 700Tavenas 1971 12-BP74 8.50 Granular Granular 446Thompson 1979 305 mm square concrete 15.10 Granular Cohesive 1560Vesic 1970 457 mm, closed-bottom pipe 15.00 Granular Granular 3200Vesic 1970 457 mm, closed-bottom pipe 12.00 Granular Granular 2630Vesic 1970 457 mm, closed-bottom pipe 8.90 Granular Granular 1872Vesic 1970 457 mm, closed-bottom pipe 6.10 Granular Granular 1533Vesic 1970 457 mm, closed-bottom pipe 3.00 Granular Granular 411Vesic 1970 406 mm square concrete 15.20 Granular Granular 2310Walkinshaw and Healow 1994 356 mm square concrete 39.60 Cohesive Cohesive 2805aWalkinshaw and Healow 1994 406 mm, closed-bottom pipe 32.20 Cohesive Cohesive 1335aWalkinshaw and Healow 1994 356 mm square concrete 32.60 Cohesive Cohesive 1160aWalkinshaw and Healow 1994 437 mm timber 32.00 Cohesive Cohesive 1335a

Table 1. Summary of the pile load tests.

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again. The 55% value was chosen because it gave the bestoverall bearing capacity predictions. It should be noted, how-ever, that any value between 40 and 60% will give the samebearing capacity prediction for the driven piles included in thedatabase.

Whenever an open-bottom pile is considered, the full toearea is normally used to compute its toe bearing capacity. Thisassumption is only true if the friction between the soil plug andthe pile inner walls is greater than the pressure applied to thesoil by the pile toe. When piles of this type get close to or startto penetrate a soil stratum resting underneath a somewhatweaker soil deposit, the true cross section of the pile should beused to compute its toe bearing capacity, even though such anapproach is conservative because it ignores the friction be-tween the soil plug and the pile.

Design of piles in granular soilsMeyerhof (1976) proposed a simple pile design method forgranular soils. It is based on the standard penetration test (SPT)N values. He used the relation proposed by Peck et al. (1974)to correct the SPT N values. Samson et al. (1986) proposed themore general correction factor CN given by

[9] CN = 95.76vThis relation is very similar to the one used by Meyerhof, butit is easier to include into a computer program. The unit weightof the soils must be known to compute the effective verticalstress v. Bowles (1977) suggested the relation between theunit weight of granular materials and the corrected SPT Nvalues shown in Fig. 1. The author found that this relationworks well enough with uncorrected SPT N values and used itto estimate the unit weight of the soils whenever it was notprovided in the papers.

As mentioned by Meyerhof (1976), the author found thatusing SPT N values which are corrected for the overburdenpressure to estimate the unit toe resistance and uncorrectedSPT N values to predict the unit skin friction will provide thebest total bearing capacity prediction.

The following equation was found to give the best overallprediction of the ultimate capacity of the test piles installed ingranular soils:[10] fs = 1.9N

It is very close to eq. [7a] proposed by Meyerhof and seemsto work equally well for displacement piles and so-called low-displacement driven piles (H piles and open-bottom pipepiles). It also seems to give good results for bored piles in sand.

It should be pointed out, however, that this conclusion is onlyvalid for single piles. The bearing capacity of piles driven in-side a large group of piles is likely to be affected by the com-paction of the soil caused by nearby piles. It should also benoted that these comments are only true for piles with uniformcross sections. Timber piles, with their conical shape, will mo-bilize a higher unit skin friction than uniform steel and con-crete piles. According to Tavenas (1971) the predicted unitskin friction of a cone-shaped timber pile should be multipliedby a factor as high as 1.8, and Meyerhof mentioned that thisvalue should be 1.5. Blanchet et al. (1980) suggested a valueof 2 for timber piles driven into a soft clay deposit. The authorfound that a value of 1.8 gave the best overall bearing capacityprediction for timber piles driven either into sand or clay. Healso used the actual pile skin surface in each soil layer to pre-dict their bearing capacity instead of an average shaft area asis often done.

The best overall prediction of the ultimate bearing capacityof the test piles installed in granular soils was given by[11a] qP = 115NPT (bored piles)[11a] qP = 190NPT (driven piles)Although eq. [11a] is quite close to eq. [6b], eq. [11b] willpredict a toe bearing capacity which is about half the valuecomputed by eq. [6a]. This large difference may be related to theuse of the Davisson failure criterion, which is often conservative

Fig. 1. Total unit weight vs. standard penetration index (Bowles1977).

Authors Pile type and dimensionsEmbeddeddepth (m)

Soil type alongthe shaft

Soil typeat the toe

Ultimatecapacity (kN)

Walkinshaw and Healow 1994 477 mm timber 32.10 Cohesive Cohesive 1650aWalkinshaw and Healow 1994 14-HP89 32.30 Cohesive Cohesive 1315aWebster et al. 1994 610 mm square concrete 13.80 Granular Granular 2200Webster et al. 1994 460 mm square concrete 27.00 Granular Granular 3200Wong 1994 500 mm bored 21.30 Cohesive Cohesive 2136Wong 1994 472 mm bored 25.90 Cohesive Cohesive 2136

a Estimated Davisson capacity.

Table 1 (concluded).

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with respect to plunging failure or other failure criteria. Itshould be noted that it was assumed that all the SPT N valueswere recorded using a donut-shaped hammer and aropecathead system which transfer a lower amount of energyto the drill rod than a true free fall driving system. SPT Nrecorded with the latter system should therefore be increasedto take into account the greater efficiency of these hammers.

Design of piles in cohesive soilsThe standard penetration test should not be used in clayey soilsunless their sand and silt contents are high enough for them tobehave as granular materials. Marine and lacustrine clays arebetter characterized by their undrained shear strength, and the

method or other similar design methods should be used todesign piles installed in clay deposits.

The Canadian foundation engineering manual (CanadianGeotechnical Society 1992) suggests the use of the valuespublished by Tomlinson (1957) to design piles installed in claywith an undrained shear strength smaller or equal to 100 kPa.Using these values gave good results but, since the clayfound at some of the sites had an undrained shear strengthgreater than 100 kPa, another relation was needed.

Coduto (1994) published the values shown in Figs. 2aand 2b for compression load tests carried out on bored anddriven piles, respectively. The curves also shown in these fig-ures are close to the arithmetic average of the data points andcan be expressed by the following equations:

[12a] = 1 (bored piles, Su 51 kPa)[12b] = 0.32 + 250Su 1.5 (bored piles, Su > 51 kPa)[13a] = 1 (driven piles, Su 32 kPa)[13b] = 0.35 + 170Su 1.6 (driven piles, Su > 32 kPa)Equations [12a][13b] were used to predict the ultimate skinfriction of the test piles installed in cohesive soil deposits and

Fig. 2. Undrained shear strength vs. for (a) bored and (b) drivenpiles (from Coduto 1994).

Pile toediameter (m)

Bearingcapacity factor

1.0 6

Table 2. Toe bearing capacity factors.

Fig. 3. Influence of pile type on calculated vs. ultimate bearingcapacity for (a) driven and (b) bored piles.

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LamVanDucHighlight

gave slightly better overall predictions than the relation suggestedin the Canadian Foundation Manual. The unit toe bearing ca-pacity was computed using equation [4] and the NC valuesrecommended in the Canadian foundation engineering man-ual (Canadian Geotechnical Society 1992). These values aresummarized in Table 2. As mentioned earlier, NC is usuallyequal to 9. The author found, however, that the lower NC valuesgave a better overall prediction for large-diameter piles.

Bearing capacity prediction

The design methods described above were used by the authorto predict the bearing capacity of the test piles included in thedatabase. Each test site was analyzed and a typical soil profilewas chosen for each of the piles. Engineering judgement wassometimes required to locate the groundwater surface or toestimate other geotechnical properties not included in the pa-pers. Any load test which required too much engineeringjudgement was discarded.

Figures 3 and 4 show the calculated bearing capacity values(Q) with respect to the actual ultimate bearing capacity (Qu)of each test pile. Figure 3 shows the influence of the pile type onthe quality of the prediction. Figure 4 shows the same results with

respect to the soil type. In this figure, sand and clay representthe main soil types. Some thin clay or sand strata may also bepresent. The term sandclay means the shaft and the toe of apile are in a different soil type. Table 3 summarizes the resultsshown in Figs. 3 and 4. In this table, the error is calculated withrespect to the actual Davisson capacity of the test piles. The errordistribution of the predicted capacity is also shown in Fig. 5.

As mentioned previously, the author only tried to match thetotal capacity of the test piles. On two test sites, however,procedures were used to insure that the test piles had no toeresistance during some of the load tests. In these cases, it istherefore possible to verify the accuracy of the predicted shaftbearing capacity.

The first of these test programs was described by ONeilland Reese (1972). Four 763 mm bored piles were installed inthe clay deposit found on the test site. One of these piles hada 2.3 m belled toe and was therefore discarded. Another pilecontained a vented void beneath the toe and was used to meas-ure the true skin friction on this pile. Table 4 gives a summaryof the characteristics of the test piles, including their ultimateand predicted capacities.

The second of these test programs was described byMcCammon and Golder (1970). Two 610 mm open-bottompipe piles were driven and tested at different depth intervals inthick granular (pile 1) and cohesive (pile 2) deposits. Pile 1was cleaned after each 12 m long section was driven. The testpile was driven to a depth of 45.4 m, a 0.3 m void was createdbeneath the pile toe to eliminate any end bearing resistancethat might increase its total capacity, and the pile was loadtested. The test pile was then closed with a 10 m concrete plug,driven to a depth of 47.2 m, and load tested again. Pile 2 was

Fig. 4. Influence of soil type on calculated vs. ultimate bearingcapacity for (a) driven and (b) bored piles.

Fig. 5. Error distribution of the predicted capacity for all test piles.

Error (%) Driven piles Bored pilesMinimum 48.5 42.3Average 2.8 8.3Maximum 40.5 12.6Standard deviation 19.3 16.2

Table 3. Statistics for predicted versus ultimate capacity.

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driven successively to depths of 30.5 and 46.6 m, cleaned outto 0.3 m ahead of the pile toe, and load tested. It was thenclosed with a 10 m concrete plug, driven to a depth of 48.2 m,and load tested again. Table 5 provides a summary of the char-acteristics of the test piles, including their ultimate and pre-dicted capacities.

Tables 4 and 5 show that there was good agreement be-tween the measured and predicted skin resistances. The totalcapacity values are also within acceptable limits.

Conclusions

The results given in Tables 35 indicate that the design proce-dures given in this paper for driven piles generally predict thebearing capacity of the test piles included in the database withan acceptable accuracy. In some cases the error with respect tothe Davisson capacity is larger than 40%, but in most casesthe predicted capacities are within 25% of the measured valuesas shown in Fig. 5.

Acknowledgements

The author would like to thank Mr. Renald Blanchet, the engi-neering director of Queformat Ltd., for reviewing this paperand providing useful comments.

References

Ahmed, S., and Sowers, G.F. 1985. Pile capacity and drivabilityacase study. Proceedings of the International Symposium on Pene-trability and Drivability of Piles, San Francisco, Calif., August1985, pp 155158.

Blanchet, R., Tavenas, F., and Garneau, R. 1980. Behaviour of fric-tion piles in soft sensitive clay. Canadian Geotechnical Journal,17: 203224.

Bowles, E.B. 1977. Foundation analysis and design. 2nd ed.McGraw-Hill Book Company, New York.

Canadian Geotechnical Society. 1992. Canadian foundation engi-neering manual. 3rd ed. Canadian Geotechnical Society. BiTechPublishers Ltd., Richmond, B.C.

Canadian Wood Council. 1991. Wood piles. Canadian Wood Council,Ottawa, Ont.

Coduto, D.P. 1994. Foundation design, principles and practices. Prentice-Hall Inc., Englewood Clifs, N.J.

Coyle, H.M., and Castello, R.R. 1981. New design correlations forpiles in sand. Journal of the Geotechnical Engineering Division,ASCE, 107(GT7): 965986

Davisson, M.T. 1972. High capacity piles. Proceedings, Soil Mechan-ics Lecture Series on Innovations in Foundation Construction,American Society of Civil Engineers, Illinois Section, Chicago1972, pp. 81112.

Endley, S.N., Ulrich, E.J., and Gray, J.B. 1979. A study of axial pileload tests. ASCE, Symposium on Deep Foundations, Atlanta, Ga.,October 1979, pp. 101121.

Federal Highway Administration (FHWA). 1991. Federal HighwayAdministration Demonstration project 66. Dynamic pile monitor-ing and pile load test report I165 1(2), Mobile County, Alabama.

Fellenius, B.H. 1994. The critical depth. How it came into being andwhy it does not exist. Background to a presentation to the 1994PDI Users Day, Orlando, Fla. Pile Dynamics Inc., Cleveland, Ohio.

Ismael, N.F., and Klym, T.W. 1979. Uplift and bearing capacity ofshort piers in sand. Journal of the Geotechnical Engineering Divi-sion, ASCE, 105(GT5): 579594

Jaime, A., and Romo, M.P. 1990. Behaviour of friction piles in Mexicoclay. Journal of the Geotechnical Engineering Division, ASCE,116: 915931.

Kraft, L.M.,Jr. 1991. Performance of axially loaded pipe piles in sand.Journal of the Geotechnical Engineering Division, ASCE, 117:272296.

Laier, J.E. 1994. Predicting the ultimate compressive capacity of along 12H74 steel pile. Proceedings, International Conferenceon Design and Construction of Deep Foundations, Orlando, Fla.,December 1994, Vol. 3, pp. 18041818.

Mansur, C.I., and Hunter, A.H. 1970. Pile testsArkansas River pro-ject. Journal of the Soil Mechanics and Foundations Division,ASCE, 96(SM5): 15451582.

McCammon, N.R., and Golder, H.Q. 1970. Some loading tests onlong pipe piles. Gotechnique, 20: 171184.

Meyerhof, G.G. 1976. Bearing capacity and settlement of pile foun-dations. Journal of the Geotechnical Engineering Division, ASCE,102(GT3): 196228.

Nordlund, R.L. 1963. Bearing capacity of piles in cohesionless soils.Journal of the Soil Mechanics and Foundations Division, ASCE,89(SM3): 1131.

ONeill, M.W., and Reese, L.C. 1972. Behaviour of bored piles inBeaumont clay. Journal of the Soil Mechanics and FoundationsDivision, ASCE, 98(SM2): 195213.

Parsons, J.D. 1966. Piling difficulties in the New York area. Journal

Pile No. Embedded depth (m) Ultimate capacity (kN) Predicted capacity (kN) Error (%) CommentsS1 7.00 1150 1382 6.9S2 14.00 2550 2801 12.6S3 7.00 1070 1015 5.1 Vented void under the toe

Table 4. Summary of the test piles described by ONeill and Reese (1972).

Pile No. Embedded depth (m) Ultimate capacity (kN) Predicted capacity (kN) Error (%) Comments1 45.4 1560 1775 13.8 Driven in sand, no toe resistance1 47.2 3967 3688 7.0 Driven in sand2 30.5 1961 1960 0.0 Driven in clay, no toe resistance2 48.2 3495 3691 5.6 Driven in clay

Table 5. Summary of the test piles described by McCammon and Golder (1970).

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of the Soil Mechanics and Foundations Division, ASCE,89(SM3): 4364.

Peck, B.P., Hanson, W.E., and Thornburn, T.H. 1974. Foundationengineering. 2nd ed. John Wiley & Sons Inc., New York.

Reese, L.C., Touma, F.T., and ONeill, M.W. 1976. Behaviour ofdrilled piers under axial loading. Journal of the Geotechnical En-gineering Division, ASCE, 102(GT3): 493510.

Samson, S.C., Liao, A.M., and Whitman, F. 1986. Overburden cor-rection factors for SPT in sand. Journal of the Geotechnical Engi-neering Division, ASCE, 112(GT3): 373377.

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Tomlinson, M.J. 1957. The adhesion of piles driven in clay soils.Proceedings, 4th International Conference on Soil Mechanics andFoundation Engineering, London, Vol. 2, pp. 6671.

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Wong, D.O. 1994. Auger-cast piles in clays. Proceedings, Interna-tional Conference on Design and Construction of Deep Founda-tions, Orlando, Fla., December 1994, Vol. 3, pp. 368384.

List of symbols

AP pile toe areaAS pile shaft area factor used to calculate the unit skin friction from

the undrained shear strength of cohesive soilc effective cohesionCN factor used to correct the SPT N values with

respect to the vertical effective stressD pile diameter friction angle between the pile shaft and the soilfS unit skin friction unit weight of the soilK horizontal earth pressure coefficientL pile lengthN standard penetration test value (blows / 0.3 m)NA average SPT N value along the pile shaftNC, N, Nq bearing capacity coefficientsNPT average SPT N value at the pile toe effective internal friction angleQ predicted ultimate capacityqP unit toe resistanceQP toe bearing capacityQS shaft bearing capacityQu ultimate bearing capacityt effective vertical stress at the pile toeV effective vertical stressSu undrained shear strength

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1997 NRC Canada

AbstractRsumIntroductionDatabasePile designBearing capacity predictionConclusionsAcknowledgementsReferencesList of symbolsTablesTable 1Table 2Table 3Table 4Table 5

FiguresFig. 1Fig. 2Fig. 3Fig. 4Fig. 5