PILE LOAD TEST FOR W.R. BENNETT BRIDGE

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1 PILE LOAD TEST FOR W.R. BENNETT BRIDGE Ernest Naesgaard, University of B.C. Ph.D. candidate & Naesgaard Geotechnical Ltd., B.C., Canada Uthaya Uthayakumar, Trow Associates Inc, B.C., Canada Turgut Ersoy, Victoria, B.C. Canada Don Gillespie, Ministry of Transportation, Victoria, B.C., Canada ABSTRACT A pile load test has been conducted in order to obtain design information for a new five lane W.R. Bennett Bridge across Okanagan Lake at Kelowna, B.C. Soils at the site consist of soft silts over loose to medium dense sandy silts and silty sands to a depth greater than 100m. The load test included installation of five 610 diameter pipe piles to a depth of approximately 45m below the lake bottom. Four of the piles were reaction piles for a central fifth axial test pile. Pile driving analyzer (PDA) monitoring and capacity analyses (CAPWAP) were preformed for all five piles. A static axial load test to failure was carried out on the central pile which was instrumented with strain gauges and tell-tales at mid-depth and bottom locations. The tests showed good agreement between measured pile capacity and PDA (CAPWAP) measurements. Calculated capacities vary significantly and demonstrate the importance of conducting a load test and/or PDA measurements if local experience with the pile and soil type is not available. Residual loads in the pile are shown to have a significant effect on the interpreted distribution of shaft and toe capacity. The test also showed that residual loads induced from static loading were much higher than those induced by impact driving. On completion of the axial load test, two of the reaction piles were tested laterally by jacking the piles apart and together for several cycles. RÉSUMÉ Un essai de charge de pieu a été effectué afin d'obtenir l'information de conception pour un nouveau W.R. Bennett Bridge à cinq ruelles à travers le lac Okanagan à Kelowna, B.C. Les sols à l'emplacement se composent de silts mou sur des silts sableux à densité lâches à moyenne et sables silteux à une profondeur de plus de 100m. L'essai de charge a inclus l'installation de cinq 610 pieux en tuyau de diamètre avec une profondeur approximativement de 45m audessous du fond de lac. Quatre pieux étaient des pieux de réaction pour un cinquième pieu central d essai. Surveillance du pieu conduisant l'analyseur (PDA) et analyses de capacité (CAPWAP) ont été préformées pour chacun des cinq pieux à la fin du fonçage initial et sur le re-frappant suivant les essais. Un essai axial statique de charge à rupture a été effectué sur le pieu centrale qui a été équipée avec des jauges de déformation et des tell-tales aux endroits de miprofondeur et de fond. Les essais ont montré de bonnes corrélations entre la capacité mesurée du pieu et les mesures de PDA (CAPWAP). Les capacités calculées changent de manière significative et démontrent l'importance d'effectuer un essai de charge et/ou mesures de PDA si une expérience locale avec le type de pieu et de sol n'est pas disponible. Les efforts résiduels dans le pieu montrent avoir un effet significatif sur la distribution interprétée de la capacité le long du fut et à la pointe. L'essai a également prouvé que les efforts résiduels induits du chargement statique étaient beaucoup plus hauts que ceux induits par la conduite d'impact. Sur l'accomplissement de l'essai axial de charge, deux des pieux de réaction ont été examinés latéralement par vérinage des pieux à part et ensemble pour plusieurs cycles. 1. INTRODUCTION The British Columbia Ministry of Transportation is replacing the existing 1.5 km long three lane Okanagan Lake Bridge with a new five lane structure (Fig.1). A combination of floating bridge, pile supported fixed structure, and light-weight fill approach structure are proposed. The bridge site has a water crossing of 920m, water depths up to 45m, and weak compressible foundation soils. The fixed structure is to be supported on 36 m to 49 m deep 610mm and 914mm diameter pipe piles. A pile load test program was conducted in April 1999 for the proposed new bridge to assess axial pile capacity, lateral load-deformation behaviour, and drivability. The test program included: cone penetration testing (CPT) (Fig. 2), driving of five 610mm diameter x 12.7mm thick steel pipe piles to 45 m below the lake bottom, pile driving analyzer (PDA) testing of all the piles at end of initial driving and on a re-strike 23 days later, instrumentation of the test pile with strain gages and tell tales with monitoring during installation and load testing, axial load testing of the central test pile, and lateral load testing of two of the reaction piles. Lateral deformation was monitored by inclinometer measurements within casings installed in the centre of each pile. The test program provided data and insight on: the affect vibratory driving on capacity, residual loads and load distribution within the pile, comparison Figure 1 Location of Pile Load test

2 Figure 2 Cone penetration test (CPT) at location of test pile of capacity of open-ended and closed-ended piles, comparison of axial capacity from PDA testing and static load testing, comparisons of the measured axial capacity to those calculated using various design methods, and comparison of lateral behaviour to that calculated using p-y procedures. Fig. 3 shows layout of the test piles in plan and profile. Water depth at the pile load test was about 6.5 m DEPTH * (m) 0-7 Table 1. Soil profile at pile load test site DESCRIPTION Very soft to firm low plastic SILT and occasional layers of organic matter and silty fine sand Moisture content - 23 to 74 %, Liquid limit - 24 to 58 %, Plastic limit - 17 to 44 %, 7-14 Sandy SILT with layers of silty fine sand Soft to firm low plastic SILT with layers of fine sand and clay. Moisture content - 23 to 59 %, Liquid limit - 23 to 65 %, Plastic limit - 18 to 31 %, Loose to medium dense silty fine SAND with some SILT layers. >44 Dense silty fine SAND with occasional SILT layers * - depth below mudline (approx. elev. 336 m geodetic) 2. GEOLOGY AND SOIL PROPERTIES The Okanagan lake occupies a relatively narrow elongated north-south trending fjord-like trench of approximately 120 km length, 3.5 km width and over 2,000m deep (Eyles et al, 1990). During the Pleistocene era the trench was scoured by several glacial advances (Fulton, 1975) (Nasmith, 1962), the last being approximately 15,000 years ago. As the ice level dropped a lake was formed in the basin into which large quantities of sands, silty sands, silt, and clay sediments were deposited. These deposits, initially glacio-lacustrine and later lacustrine, drape over the underlying topography. Based on the geological history the sediments in the lake are expected to be normally consolidated and at the site of the pile load test were over 60 m in depth. The sub-soil in the vicinity of the pile load test is summarized in Table 1 and Fig PILE INSTALLATION Piles were installed from a 80m by 15m barge, using an 80 tonne crawler crane. The five pipe piles (P1 to P5 in Fig. 3) were supplied in 18.3m lengths and spliced on site. Three hammers were used to install the piles; a 3,855 kg drop hammer, a 5,445 kg drop hammer, and an APE vibratory hammer. The initial sections of three piles were driven with the vibratory hammer to check if vibratory driving had any adverse affect on pile capacity. A Delmag D62 diesel hammer was used to re-strike the piles eight days after

3 the axial load test. Details of the pile installation are summarized on Table 2 and Fig PILE DRIVING ANALYZER TESTING Pile driving analyzer (PDA) testing was conducted at end of initial driving (EOID) and 23 days after the EOID (8 days after the axial pile load test). Small and large drop hammers were used for EOID testing and large drop hammer and Delmag D62 diesel hammer were used for the final re-strike testing. CAPWAP analyses were conducted to determine pile capacities from the PDA (see Table 3). Residual or locked-in loads were not considered in PDA/CAPWAP shaft and toe capacities in Table 3. Figure 4 Driving record of pile P1. The final set with PDA at 10 am on Mar. 24 was 8 blows/25mm, on Mar. 24 at 3:45 pm with 5,445 kg hammer set at 13 blows/25mm & on re-strike on Apr. 16 with Delmag D62 set at 19 blows/25mm. Pile was driven 400 mm with D62 on April AXIAL PILE LOAD TEST 5.1 Axial Pile Load test Instrumentation Figure 3 Axial pile load test arrangement The test pile (P1) was instrumented with vibrating wire strain gages (four per level) and tell-tales (Fig. 3) prior to driving of the pile so residual loads from pile installation and load distribution during the test could be monitored. Pick-ups on two lower strain gages dislodged during vibratory driving, but were restored without loss of data. Pile settlement was monitored using six displacement transducers (LVDTs); four on the perimeter of the pile head and one on each tell-tale. The LVDTs were attached to a steel reference frame that was independent of the test piles and supported by four 324mm diameter pipe piles. Pile elevations were also measured during the load test by sighting with a transit on scales attached to the test and reaction piles. The transit was set on shore 18 m from the test piles. This redundancy of data proved to be useful as the LVDTs reached the end of their travel several times during the test and had to be reset. Axial load applied to the pile was measured using two calibrated 2.7 MN

4 Time relative to EOID (days) Table 2 - Summary of Pile installation & testing sequence P1 P2 P3 P4 P5 CLOSED-TOE OPEN-TOE CLOSED-TOE CLOSED-TOE OPEN-TOE -8 - D S (0-8.2) - D S (0-9.1) D S(0-9.1) S(18.9) V( ) (20min) S(18.0) V( ) (1.2min) D S(0-9.1) S(18.9) - S(18.9) V( ) (30min) -5 D S( ) - D S( D S( ) - D S(0-90.1) -2 - S(37.2) S(37.2) S(37.2) S(37.8) -1 S(38.2) D S( ) D S( ) D S( ) D S( ) - EOID D S( ) PDA w/ D S & D L D S( ) PDA w/ D L D S( ) PDA w/ D L D L( ) PDA w/ D L D L( ) PDA w/ D L CONCRETE FILL CONCRETE FILL - 15 AXIAL LOAD TEST / LATERAL LOAD TEST LATERAL LOAD TEST - 23 re-strike PDA w/ D 62 PDA w/ D 62 PDA w/ D L PDA w/ D L PDA w/ D 62 Toe elev.(m) Legend for Table 2 D S = 3855kg drop hammer with 2.5m drop; D L = 5445kg drop hammer with 2.5m drop; V = APE vibratory hammer; D 62 = Delmag D62 diesel hammer; ( ) = driven from 21.0 to 28.4 m depth with indicated hammer; S(18.9) = Splice at 18.9m from toe of pile; (20min) = Duration of vibratory driving in minutes; EOID = End Of Initial Driving; PDA = Pile Driver Analyzer testing load cells in parallel on the pile head. Jack pressure was monitored; however the 7.1 MN hydraulic jack was not calibrated. Electronic data collection was done with Model C10 RST data logger and laptop computer. 5.2 Axial Pile Load Test The test was conducted on the central pile P1 in three stages (Fig. 5 & 6). The first stage consisted of a standard load and unload cycle in general accordance with the "quick test" procedure of ASTM D1143. The pile was loaded in increments to a failure load of 3,750 kn. Each load increment was 15 minutes. The second Figure 5 Load vs deformation at head of pile for the axial test on pile P1 stage consisted of 21 cycles of load oscillation between 2,000 kn and 3,000 kn. The cyclic frequency was governed by the jack's loading speed and was about five minutes per cycle. The third stage consisted of a final loading to failure followed by unloading with quick increments. The maximum load attained was 4,000 kn, however this load could not be sustained due to creep settlement of the pile. A summary of axial load capacity from the test, from PDA measurements and various calculation methods is in Table 3. Failure loads from the load test in Table 3 were derived using the approach by Davisson (CFEM, 1992). 5.3 Calculated Axial Capacity Axial pile capacities were calculated using direct CPT methods as described by Eslami & Fellenius, 1997, and by the API RP2A, 1987 method as listed below (bold letters refer to titles in Table 3): French or LCPC method (LCPC)) Eslami-Fellenius method (E-F) European method (D) Meyerhof method (M) & Schmertmann method (S) API, 1987 (API) The program UNICONE (Fellenius & Infante, 2002) was used for all the above procedures except the LCPC (Bustamante and Gianeselli, 1982) and API methods which were carried out using spreadsheet calculations. In the LCPC, European, and Meyerhof methods both shaft and toe capacities are a function of

5 Figure 6 Load & settlement versus time for axial load test on pile P1 (Apr. 16). Toe readings are from gages 1 m above the toe and middle readings are from gages 17.5 m above the toe. Letters in brackets refer to time of inferred load distributions shown in Fig. 7. Toe and middle settlements were calculated from measured stresses, pile stiffness and pile head settlement as tell tale readings were inconsistent. CPT tip bearing (q c). The Meyerhof method is only recommended for use in non-cohesive soils. In the modified LCPC method the soil type was determined using the program UBCINT (v.5.2) and the capacity was calculated using corrected CPT tip bearing (q t) instead of q c. In the Schertmann method the shaft capacity is calculated as a function of CPT sleeve friction (f s) and the toe capacity as a function of q c. The Eslami & Fellenius method shaft and toe capacity are a function of an effective CPT tip resistance (q e = q t u) where u is the CPT pore pressure measurement. Capacities from various procedures are summarized on Table 3. Effective stress β and N t parameters (CFEM, 1992), (Fellenius, 1999) have also been backcalculated from the load test data as follows: Q s = π D β σ v ' s & Q t = π (D 2 /4 ) N t σ v ' t or β = 2 Q s / (π D γ' z 2 ) & N t = 4 Q t /(π (D 2 γ' z) Figure 7 Load distribution in pile shaft inferred from strain gages & top load. Strain gages were zeroed prior to pile installation. (a) is start of load test, (b) is initial peak load, (c) is first unload, (d) is final loading, (e) is final unloading (similar to (c), & (f) is re-strike 8 days later. Note that measurements were only made at top of pile and at the gages, the load distribution between may vary from that shown. where β = shaft resistance factor N t = toe bearing capacity factor = summation over pile length D = pile diameter σ v ' s = vertical effective stress along shaft σ v ' t = vertical effective stress at toe Q s = total shaft capacity Q t = toe capacity

6 Case Pile shaft capacity (kn) Pile toe capacity (kn) Total capacity (shaft and toe) Pile No. EOID (day 0) Table 3 Summary of Axial Capacities of piles in MN PDA testing Calculated Capacity Pile Load Test (1) (day 15) Restrike (day 23) LCPC E-F D M S API post-driving residual stresses ignored included P to to to 1.5 P2 3.3 to P P4 2.6 to to 3.5 P5 2.6 to P to to to 2.5 P2 0.4 to P P to 0.9 P5 0.3 to P to to to 4.0 P2 3.8 to P P4 3.0 to to 4.4 (kn) P5 3.0 to Notes: (1) Shaft and toe load distribution has been corrected by subtracting 110 kn to compensate for lower strain gauges being one meter above the pile toe. Piles P1, P3, & P4 were closed-toed and P2 and P5 were open-toed. (2) Calculated capacities will vary according to parameters chosen and are not given to show that one method is better than another but rather to show the variability in commonly used design methods. γ' = buoyant unit weight z = pile embedment depth If residual loads in the pile are considered, then β is 0.09 and N t is 22. If residual stresses are ignored, β would be 0.15 and N t would be 15. Clearly knowledge of residual loads in the pile has a significant influence on the back-calculated parameters. 5.4 Discussion on Axial Capacity and Load Test Pile Installation Considerations: The 610 mm diameter pipe piles were installed using both vibratory and impact driving and with open and closed toes. Piles P1, P2 and P4 were vibratory driven from about 9 m below the mudline to about 25 m below the mudline (see Table 1). In all cases the piles were impact driven for the remainder depth. The PDA derived capacities for the piles with initial sections vibratory driven had similar capacities to pile P3 and P5 that were only impact driven only. Similarly piles P1, P3, and P4 with closed-toes had similar capacities to Piles P2 and P5 which had open-toes. The work demonstrated that vibratory driving of closed-toed piles in these soils is not efficient and not recommended. The closed-toed piles were slow to advance, bounced excessively, and suffered fatigue failures at the clamps when vibratory driven. Residual or locked-in load: The importance of residual loads in assessing the correct load distribution in piles was noted as early as 1963 by Nordlund. A historical review is given by Mouta da Costa et al and the concepts are discussed in detail by Fellenius, 2002a & 2002b. Residual loads can be induced in the pile by driving or jacking and by post-driving consolidation settlement of the soil. At this site the main source is believed to be from the driving. Driving impact on a pile causes a compression wave to travel down the pile. Following passage of the wave the pile attempts to rebound to its former stress state, however the soil around the pile prevents full rebound and induces residual loads in the pile. The axial load test at this site clearly demonstrated the importance of residual loads in interpreting load distribution in the pile. As indicated in Fig. 6 and 7, pile installation induced a residual load of 835 kn near the pile toe and 400 kn 17.5 m above the toe to give an approximate load distribution (a) in Fig. 7. Applying load to the pile head compresses the pile shaft and initially reduces the shear stresses along the side of the pile. It is interesting to note that at a time of approximately 14:30, (similar situations exist at times 18:30, 20:30 & 23:15) in Fig. 6 there is almost no shear stress on the pile shaft and nearly all the applied load is resisted by the pile toe. Continued loading after 14:30 causes opposite shear stresses to build up along the pile shaft until (b) in Fig. 7 when the pile is near its capacity. At (c) and (e) in Fig. 7 the load on the pile head is zero and the residual stresses near the pile toe are almost double what they were at the start of the test. This is indicative of the residual stresses from static pushing being higher than that from impact driving. When the pile was re-struck 8 days after the load test ((f) in Fig. 7) the residual stresses reduced to values similar to those at the start of test. When residual stresses are considered approximately 60% of the capacity of pile P1 is taken by the toe and only 40% by the shaft. If the strain gages had been zeroed at the start of the test (residual stresses not considered) then incorrectly 40% of the load would be attributed to the toe and 60% to the shaft. This also reflects in design parameters that may be back-calculated from the load

7 test. When residual stresses are considered an effective stress β of 0.10 and N t of 22 is backcalculated, however if initial residual stresses are ignored the values would be 0.15 and 9. Clearly extrapolation of the load test results to other lengths and diameters may lead to incorrect results if the residual stresses are not considered. Cyclic Loading: Cyclically loading pile P1 between a load of 2,000 kn and 3,000 kn (about 3/4 of the pile capacity) resulted in almost no oscillation in load at the pile toe and about 1/2 the applied amplitude at the midlevel strain gages. It is postulated that, for the pile length to stiffness ratios similar to test pile, load oscillations that do not exceed the geotechnical shaft capacity will not transfer to the pile toe. Increase in Pile Capacity: During the load test, pile P1 was initially loaded to failure at approximately 3,600 kn, then the pile was unloaded and subsequently reloaded to failure at approximately 3,950 kn. The increase in capacity is nearly all from the toe ((b) and (d) in Fig. 7 are nearly parallel) and may be due to the pile toe moving deeper into the strong sandy layer at approximately 45m depth. Comparison of Pile Capacities: In this discussion it is assumed that the correct pile capacity is from the static load test with residual or locked-in load accounted for. With respect to total pile capacities the agreement between the PDA "CAPWAP" capacity (at re-strike), the CPT modified LCPC capacity and Meyerhof method capacity, and that from the pile load test (3,950 kn) is good (within ±15%). In all cases the load test capacity was lower than those calculated and slightly lower than the PDA/CAPWAP results. The PDA/CAPWAP on redrive gave results which were only 5 to 8% higher than the pile load test capacity of 3,950 kn, the modified LCPC method gave capacity which is 8% higher, and the Meyerhof method gave a capacity that is 14% higher. When the loading is broken down to shaft and toe distribution the agreement is not as good. All methods considered were within ±40% of the toe capacity with the Eslami and Fellenius method and Meyerhof method giving very good predictions. The scatter in the calculated shaft capacities was much larger and all methods over-predicted shaft capacity. The Eslami & Fellenius, Schmertmann, European and API methods all over-predicted shaft capacity by more than 100%. Overall the best prediction was by the Meyerhof method and the worst was the European method. The back-calculated effective stress β of 0.10 and N t of 22 are generally lower than commonly quoted values (CFEM, 1992). Figure 8 Elevation showing configuration of the lateral load test of the laterally loaded piles, jack and instrumentation is shown in Fig mm diameter inclinometer casings were installed centrally in closed-ended and concrete filled piles P3 and P4. During the lateral test inclinometer measurements were made at each load increment. Lateral travel of the pile head relative to the reaction frame was measured using both an electronic wire extensometer and a tape measure. Lateral load was calculated from the jack pressure using a calibrated gage. The comparison between the calculated capacities and load test results clearly demonstrates the importance of load testing and/or PDA testing piles. 6. LATERAL PILE LOAD TESTING Only a brief overview of the lateral load test is given in this paper due to space limitations. The configuration Figure 9 Displacement profiles for Pile P3 at various lateral loads

8 The load test was conducted with the standard procedures of ASTM D3966 with minor modifications. The test piles were loaded using a 300 kn hydraulic ram with a travel of 1,200 mm. First the piles were loaded in increments to a jack ram displacement of 1,155 mm and a load of 146 kn. The time interval between each load increment was about 30 minutes to allow time for inclinometer readings in each pile. Following this loading and unloading the piles were cycled between their original location and maximum displacement. Fig. 9 shows the displaced profiles of pile P3 with various first stage loadings. The behaviour of pile P4 was similar. Reversal of the applied load indicated that there was friction in the loading frame up to about 20 kn. When this friction was accounted for the deflected shapes of the pile approximately matched that obtained using p-y methods (ENSOFT, 1996). 7. CONCLUSIONS A pile load test in the very loose to medium dense silt to fine sand soils of Okanagan Lake in central British Columbia demonstrated the importance of conducting a pile load test and/or pile driving analyses (PDA) when installing piles in unfamiliar soils. Conclusions from the tests were: (i) Open and closed-toed piles (similar to the tested piles) have similar capacity, (ii) Vibratory driving of the upper half of the pile did not affect capacity, (iii) PDA/CAPWAP and the static load test gave similar capacities, (iv) Calculated pile capacities can vary widely from actual test capacities and should not be relied upon without calibrating with load test, PDA analyses or other local experience, (v) Residual loads from pile installation are significant and knowledge of them is required to get the correct load distribution from the test. (vi) The installation method can have a significant effect on residual loads in the pile. Residual loads due to static jacking were almost double those due to impact driving. (vii) Cyclic loading between half and three quarters of the failure load did not cause significant load oscillation at the test pile toe. It is postulated that this is the case as long as the pile shaft capacity is greater than the oscillation amplitude. A description and some results from a lateral load test carried out as part of the test program is also given but not discussed in detail. 8. ACKNOWLEDGEMENTS Westmar Consultants Inc. was the prime consultant, Macleod Geotechnical Ltd. (Trow Associates Inc.) the geotechnical consultant, and Griffiths Pile Driving the contractor to the Ministry of Transportation for the pile load test. The authors thank Malcolm MacFaden and Peter Peart for their valiant and successful repair of the two strain gages at the bottom of the 55m deep pipe pile, Dean Polvi of RST Instruments Ltd. for instrumentation, Pascale Rouse for abstract translation, Bengt Fellenius for review of the paper and Victor Szabo for supervising the pile load tests and compiling data. We also thank the Ministry of Transportation, Westmar Consultants Inc., Trow Associates Inc., and SNC Lavalin Inc. for permission to publish this work. References ASTM D1143, 1981, Standard Test Method for Piles under Static Axial Compressive Load, ASTM, Philadelphia, Pennsylvania, USA. ASTM D3966, 1995, Standard Test Method for Piles under Lateral Load, ASTM, Philadelphia, Pennsylvania, USA. API, RP 2A-WSD, Recommended Practice for Planning, Designing and Constructing Fixed Offshore Platforms - Working Stress Design, 20 th edition, American Petroleum Institute, 1220 L Street, Northwest Washington, DC 20005, July 01. CFEM, Canadian Foundation Engineering Manual, 3 rd Edition, Canadian Geotechnical Society c/o BiTech Publishers, Richmond, B.C. Ensoft, Program GROUP for the Analysis of a Goup of Piles subjected to axial and lateral loading, Ensoft Inc. Austin, Texas. Eslami, A. and Fellenius, B.H., Pile capacity by direct CPT and CPTu methods applied to 102 case histories, Can. Geot. J., 34(6), pp , Dec. Eyles N., Mullins, H., and Hine, A.C., Thick and fast: Sedimentation in a Pleistocene fjord lake of British Columbia, Canada, Geology, Vol. 18, pp , Nov. Fulton, R.J., Quaternary geology and geomorphology, Nicola-Vernon area, British Columbia, Geological Survey of Canada, Memoir 380. Fellenius, B.H., Basics of Foundation Design, 2 nd Edition, BiTech Publishers, Richmond, B.C., 164 p. Fellenius, B.H., and Infante, J-A, UNICONE USER MANUAL version 1, Unisoft Ltd., 1905 Alexander St. SE, Calgary, Alberta, T2G 4J3. Fellenius, B.H., 2002a. Determining the resistance distribution in piles Part 1. Notes on shift of no-load reading and residual load, Geotechnical News, 20(2) pp , June Fellenius, B.H., 2002b. Determining the resistance distribution in piles Part 2. Method for determining the residual load, Geotechnical News, 20(3) pp , Sept. Mouta da Costa, L., Danziger, B.R., and Lopes, F.R., Prediction of residual driving stresses in piles, Can. Geotech. J. 38: pp Nasmith, H., Late glacial history and surficial deposits of the Okanagan Valley, British Columbia, BC Department of Mines and Petroleum Resources, Bulletin No. 46. Nordlund, R.L., Bearing capacity of piles in cohesionless soils, ASCE, Journal of Soil Mechanics and Foundation Engineering, 89, SM3, pp