The current drilled shaft (also called bored pile) foundation design procedures recommended in two commonly used North American foundation engineering manuals have been reviewed, and the recommended design approache from each manual is evaluated against the recent load test data conducted on continuous flight auger (CFA), cast-in-place concrete piles (augercast piles). The soil conditions where pile load tests were carried out is typical of glacial till encountered in the Canadian Prairies. The conclusion is that pile capacity prediction methods widely used in North America generally under estimate both skin resistance and end bearing for drilled shaft in very stiff to hard glacial till. For design purpose, for drilled, cast in-place concrete piles installed in glacial till soils in Western Canada, procedure recommended by Federal Highway Administration (FHWA) is recommended.

1. DRILLED SHAFT CAPACITY IN COMPRESSION – COMPARISON OF PREDICTION METHODS
Masud Karim, Ph.D., P.Eng., Cory Zubrowski, P.Eng., D. Chad LePoudre, P.Eng. SNC-Lavalin Inc., Calgary, Alberta, Canada and SNC-Lavalin Inc., Saskatoon, Saskatchewan, Canada
ABSTRACT
The current drilled shaft (also called bored pile) foundation design procedures recommended in two commonly used North American foundation engineering manuals have been reviewed, and the recommended design approache from each manual is evaluated against the recent load test data conducted on continuous flight auger (CFA), cast-in-place concrete piles (augercast piles). The soil conditions where pile load tests were carried out is typical of glacial till encountered in the Canadian Prairies. The conclusion is that pile capacity prediction methods widely used in North America generally under estimate both skin resistance and end bearing for drilled shaft in very stiff to hard glacial till. For design purpose, for drilled, cast in-place concrete piles installed in glacial till soils in Western Canada, procedure recommended by Federal Highway Administration (FHWA) is recommended.
RÉSUMÉ
Après un examen des procédures actuelles de conception de fondations sur pieux forés (aussi appelés « pieux forés moulés dans le sol ») recommandées par deux manuels d’ingénierie des fondations communément utilisés en Amérique du Nord, la méthode de conception recommandée par chaque manuel a été évaluée en fonction des données d’essais de charge menés récemment sur des pieux moulés dans le sol à l’aide d’une tarière continue (CFA), aussi appelés « pieux forés à la tarière continue ». Ces essais de charge des pieux ont été réalisés sur un terrain caractéristique du till présent dans les Prairies canadiennes. Il a été conclu qu’en règle générale, les méthodes de prévision de la capacité des pieux couramment utilisées en Amérique du Nord sous-estiment la résistance de frottement et de pointe des puits forés dans le till très ferme à dur. Pour la conception de pieux forés moulés dans le sol installés dans le till de l’Ouest canadien, il est recommandé d’utiliser la procédure prescrite par la Federal Highway Administration (FHA).

2. 1 INTRODUCTION
Most pile text books and engineering manuals discussed two basic approaches for pile capacity estimation using static analysis, namely total stress (α- method) and effective stress (β-method). However, differences exist among foundation engineering practitioners as well as within foundation design manuals about the appropriate values of α and β to use for determination of shaft resistance (skin friction). Similarly, differences also exist as to the appropriate bearing capacity factor to use for estimating end (toe) bearing capacity. These discrepancies and inconsistencies are sometimes confusing to practicing foundation engineers, particularly young professionals. Also, there is no consistent method or guideline as to which method to follow in the case of intermediate soil types, such as glacial till (which is a predominant soil type in the Canadian Prairies).
In this paper, the current drilled shaft (also called bored pile) foundation design procedures recommended in two commonly used North American foundation engineering manuals will be reviewed, and the recommended design approaches from each manual will be evaluated against the recent load test data conducted on continuous flight auger (CFA), cast- in-place concrete piles. CFA piles, also known as augercast piles, are formed by drilling a continuous flight hollow stem auger into the ground, followed by pressure injection of concrete and simultaneous extraction of the auger. The sides of the hole are supported at all times by the auger, eliminating the need for temporary casing or bentonite slurry. CFA piles have been widely used in many areas within the developed world for many years, and are becoming increasingly more common in Western Canada, particularly as advances in CFA equipment and technology have occurred.
The pile load test site is located approximately 150 km to the southeast of Saskatoon, Saskatchewan. Due to a confidentiality agreement, specific details pertaining to the pile load test information will be withheld. The following reference manuals for pile design have been reviewed for the purposes of this paper:
1. Canadian Foundation Engineering Manual (CFEM), 4th Edition (2006)
2. Federal Highway Administration (FHWA), publication on Drilled Shaft Foundations (Publication No. FHWA-NHI-10-016, May 2010)
In the following paper, no distinction will be made between conventional drilled shafts and CFA piles, as basic concepts of pile capacity estimation applies to both drilled shaft and CFA pile. However, it is noted that the capacity of CFA piles generally lies between that of a drilled shaft and a driven pile (FHWA, 2010).
The main objective of this paper is to review recent load test data conducted on CFA piles (as compared to the most commonly utilized North American foundation design manuals), and to suggest a unified approach to evaluate drilled shaft pile capacity, particularly for cohesive soil types encountered in the Canadian Prairies (east of the Rocky Mountains). As an introduction, some of the basic concepts of pile capacity estimation (single piles) will be briefly discussed.
2 ESTIMATION OF PILE CAPACITY TO COMPRESSIVE LOADS
Soil-pile interaction is complex and depends on such factors as soil type, types of loads and pile installation methods (Prakash & Sharma, 1990). Therefore, pile capacity can be determined only approximately and it is highly recommended that the predicted capacity be calibrated by pile load test results, especially for large projects. Another important note is that although the total pile capacity is the sum of the capacity along the pile shaft and at pile toe, full mobilization of ultimate shaft and toe capacities develop at different relative pile-soil displacement. The pile displacement required to develop full shaft capacity is relatively small (typically less than about +10 mm), whereas the pile displacement required to mobilize full toe capacity is typically greater (typically 5 percent (%) of pile diameter for cohesive soils, but can be much higher for piles in cohesionless soils). Therefore, in most cases the working load at the base of the pile is limited by consideration of settlement rather than ultimate capacity (Terzaghi et al, 1996) and it is highly probable that in the usual range of working loads, shaft resistance is the principal load-carrying mechanism (Bowles, 1996), especially for straight shafted piles (an exception may be relatively short piles with enlarged bases (ie, belled piles)).
2.1 Total Stress Approach
For pile design purposes, total stress analysis is applicable to cohesive soils that exhibit undrained behavior under loading. The ultimate or unit resistance at a depth z along the pile shaft (fs, called skin or side or shaft friction or resistance) is determined by the undrained shear strength of soil, su, multiplied by an adhesion factor (empirical factor originally proposed by Tomlinson, 1957), α, which is mainly a function of su. The term shaft friction will be used hereafter for the pile shaft component of the pile capacity. The alpha method is considered to be a semi-empirical approach as compared to the more fundamental theoretical approach based on effective stress concept (effective stress approach). However, it is noted that the alpha method is extensively utilized and has been in use for a long period of time.
The ultimate or unit toe resistance (fb, also called end bearing) at the base of the pile is based on bearing capacity theory and is determined by su (φ’ = 0, c = su) at the base of the pile multiplied by a bearing capacity factor, Nt.
2.2 Effective Stress Approach

3. The effective stress approach, though theoretically applicable to all soil types, is generally applicable (for pile design purposes) to cohesionless soils showing drained behavior under loading or to heavily over consolidated cohesive soils, for which the long-term condition may be critical. The ultimate resistance along the pile shaft is the frictional resistance developed at the pile-soil interface (i.e., σ’h tan δ or σ’v K tan δ or σ’v β, where δ is the effective stress friction angle for pile-soil interface, K is the coefficient of lateral earth pressure (= σ’h / σ’v) and β is the shaft friction coefficient). The ultimate toe resistance at the base of the pile is taken as the ultimate bearing capacity of the soil and is determined by σ’v at the base of the pile multiplied by a bearing capacity factor, Nt. The factor Nt is same as the factor Nq and σ’v is γ’ Df more commonly used in bearing capacity theories, where γ’ is the effective unit weight of soil and Df is the depth of foundation.
2.3 α, β, and Nt Parameters – Background Information and Range
The adhesion factor, α, for shafts in soft clay and plastic silt has been found, on the basis of full-scale loading tests, to be approximately equal to su and in stiffer clays decreases with increasing strength of clay (Terzaghi et al, 1996). As explained by Tomlinson and Woodward (2009), the effect of drilling is always to cause softening of the clay along the shaft due to relief of lateral pressure on the walls of the hole. After placing concrete in the pile borehole, water migrates from the unset concrete into the clay, causing further softening of the soil. This results in lower α values for drilled shafts as compared to driven piles. The softening effect is severe for fissured clays. As discussed by Tomlinson and Woodward (2009), in clays other than London Clay, where there is no information from loading tests or publications, the adhesion factors developed by Weltman and Healey (1979, discussed later), for drilled shafts in glacial till can be used as a guide to pile design. It should be noted that the adhesion curve by Weltman and Healey were derived from data showing considerable scatter.
As discussed by Fleming et al (2008), the value of α deduced from pile load tests appears to reduce from unity or more for piles in low strength clay, down to 0.5 or less for clay with a undrained shear strength (su) greater than 100 kPa. However, due to wide scatter in data of correlation of α values with undrained shear strength, Fleming et al (2008) recommended either α, as a function of the strength ratio, su/σ’v (based on research by Randolph and Wroth, 1982) or effective stress approach using β values for piles in clay.
CFEM suggested use of correlation between α and su developed by Stas & Kulhawy (1984) for drilled shafts based on both uplift and compression test data. FHWA (2010) recommended a constant α value for su/pa less than 1.5, where pa is the atmospheric pressure in the same units as su and following equation is for 1.5 ≤ su/pa ≤ 2.5.
훼 =0.55−0.1൜ 푠푢 푝푎 −1.5ൠ
No definitive recommendations about proper β values could be found in Tomlinson and Woodward (2009) and Fleming et al (2008). As discussed in the FHWA Drilled Shaft Reference Manual (2010), various design models have been proposed for evaluating the β term in effective stress approach. The method currently recommended in AASHTO (2007) is the “O’Neill and Reese (1999)” method, where β is calculated solely as a function of depth below the ground surface, without explicit consideration of soil strength or the in-situ state of stress. This approach is based on fitting a design curve to values of back-calculated β from field load tests. A more rational approach, as presented for example by Chen and Kulhawy (2002) and Kulhawy and Chen (2007), is to evaluate separately values of K and δ which are then combined to determine β. Results of research published over the past 15 years demonstrate that this approach can provide reliable estimates of side resistance and represents a rational method to incorporate soil strength and state of stress into design equations. FHWA (2010), therefore, recommends that designers employ this model. In this method, β for cohesionless soils can be approximated by:
훽 ≈ (1− sin∅′)ቊ 휎′푝 휎′푣 ቋ sin∅′ tan∅′≤퐾푝tan∅′
Where σ’p is the effective pre-consolidation pressure and is estimated by the following equation:
For sandy soils:
휎′푝 푝푎 ≈0.47 (푁60)푚
where:
m = 0.6 for clean quartzitic sands and m = 0.8 for silty sands to sandy silts
pa = atmospheric pressure in the same units as σ’p (eg, 2,116 psf or 101.3 kPa)
For gravelly soils,
휎′ 푝 푝푎 =0.15 푁60
For cohesive soils, the bearing capacity factor Nt usually varies between six and nine and for cohesionless soils, Nt varies with soil’s internal friction angle, φ. For all soil types, both Tomlinson and

4. Woodward (2009) and Fleming et al (2008) recommended Nt of value of 9 based on work by Skempton (1951) provided that the pile has been installed at least to a depth of five diameters (Tomlinson & Woodward, 2009) or three diameters (Fleming et al, 2008) into the bearing stratum. This is supported by Terzaghi et al (1996). For cohesive soils, FHWA (2010) recommended an Nt value of 9 only for cases where the shaft depth is at least three times the diameter and the mean su is about 96 kPa (2,000 psf). For smaller values of su, Nt can be approximated as a function of su as given in Table 1. Linear interpolation can be used for values between those tabulated. On the other hand, CFEM recommended Nt values based on pile diameter, also shown in Table 1.
Table 1. Bearing capacity factor, Nt
2.4 Approach for Soils other than Pure Cohesionless and Cohesive Soils
Glacial till is the most dominant soil type in the Canadian Prairies. Glacial till is a heterogeneous mixture of sand, silt, clay and gravel, and typically contains cobbles and boulders. As such, glacial till exhibits unique characteristics that differ from purely cohesionless soils (ie, gravel, sand and silt) and purely cohesive soils (clay). The majority of glacial till is found to be low to medium plasticity. Although glacial till is typically considered to be a good foundation material because of relatively high strength and low compressibility, especially till encountered at depths, it also poses difficulties for pile construction due to the presence of cobbles and boulder, as well as sorted granular deposits (which are often seepage/sloughing zones).
Glacial till is often interbedded with other types of (sorted) glacial deposits, and can show characteristics intermediate between cohesionless and cohesive soil. Therefore, till may behave differently than purely undrained soil immediately after pile installation. As such, prediction methods based on pure cohesive behavior may not be applicable for piles in all glacial till deposits. In FHWA (2010), glacial till was listed as one of the geomaterials requiring special consideration.
3 ESTIMATION OF PILE COMPRESSION CAPACITY USING EXISTING GUIDELINES
Following is a worked example of estimated pile capacity using geotechnical data from near the recent CFA pile load test locations. The recommended procedures outlined in the above mentioned manuals were followed. In addition, the adhesion factors for piles in glacial soils developed by Weltman and Healey was used for comparison purpose. The individual shaft and toe capacities will be compared with the measured capacities from the recent pile load test data, as described below.
3.1 Soil Conditions
The general stratigraphy (up to 35 m maximum depth drilled) consisted of glacial till, (typically silt and sand, some clay, trace to some gravel) overlain by surficial deposits of sand and silt. Numerous sand seams were observed at various depths, especially below 20 m depth. Wet sand layers were encountered at depths of 17.2 and 19.8 m at a borehole near the pile load test area. Discontinuous gravel deposits were encountered in some boreholes. Cobbles and boulders were encountered in some of the boreholes and in the zone from about 5.1 to 9.5 m, and then in the zone from about 21.7 to 30.8 m near to the pile load test area.
Based on the SPT ‘N’ values, the till was typically hard in consistency. However, two softer (stiff to very stiff) layers were encountered in some boreholes: one stiff and relatively thin, this is immediately below surficial deposits at elevations of about 540 to 538 m (1 to 4 m below existing ground surface), and two stiff to very stiff, starting at an elevation of about 533 to 532 m (9 to 10 m below existing ground), about 3 to 4m in thickness.
SPT N values vs elevations are plotted in Figure 1. Figure 2 shows the average N value profile, which indicates that N values initially, increases with depth, then decrease (softer zone) at around 9 to 10m below ground and increase again at around 12 to 13 m.
Unconfined compression strength (UCS) testing was conducted on selected glacial till samples. In addition, three unconsolidated undrained (UU) Triaxial tests were performed on till samples (it is noted that more UU tests were planned, but that recovering intact till samples was difficult due to the presence of cobbles/boulders and the hard nature of the till deposit with depth). The results are summarized in Table 2.
Table 2. Summary of UU triaxial testing
Sample Number
Depth (m)
Su1 (kPa)
JMT-112
6.1 – 6.5
260
JMT-131
18.3 – 18.7
325
JMT-151
30.5 – 30.9
295
1 Undrained shear strength.
During drilling of borehole 506309-001, SPTs were conducted near the depths at which Shelby tube samples were recovered. Using the SPT ‘N’ value closest to the Triaxial samples, a Su/N ratio of 9.5, 5.8 and 6.3 was obtained for samples JMT-112, 131 and 151 respectively. This confirmed the typical Su/N ratio of 6.0 to 6.5, except one case where the ratio is unusually high (sample JMT-112). Upon review of the
Recommending Agency
su (kPa)
Pile Diameter (m)
Nt
FHWA
25
-
6.5
50
8.0
100
9.0
CFEM
-
< 0.5
9.0
0.5 – 1.0
7.0
> 1.0
6.0

5. sample on completion of the UU test, it became apparent that the test result was influenced by the presence of a relatively large rock (coarse gravel) near the centre of the sample. Based on a correlation of Su/N of 6.1, the probable (estimated) undrained shear strength of sample JMT-122 sample is ~160 kPa.
Based on the above observations, Su/N equal to 6.1 was adopted to obtain Su profile with depth.
Figure 1. Uncorrected SPT blow counts (N value) vs elevation
Figure 2. Average uncorrected ‘N’ value profile
3.2 Groundwater Conditions
Only one standpipe piezometer was located in the vicinity of the test piles and the nearest boreholes. Two more standpipes were installed further east and south of the nearest borehole location. Based on the piezometers readings, the shallow groundwater elevations are very close to the existing ground level (0 to 2 m below existing ground) varied from 541 to 539 m. It should be noted that screen levels varies from 537 to 534 m elevations (about 4 to 7 m below existing ground). In addition, four vibrating wire piezometers were installed in one of the boreholes. The piezometers were installed at the design depths of 10, 15, 25 and 35 m below ground surface. The piezometers were grouted into the borehole with a cement bentonite grout. A tremie pipe was utilized to install the vibrating wire piezometers (piezometers were taped to the tremie pipe). The readings from VW piezometers at borehole 506309-001 are summarized in Table 3.
Table 3. Summary of VW piezometer readings
Date of Reading
Water Level (m)
Tip at 10
Tip at 15
Tip at 25
Tip at 35
27-Oct-12
2.061
1.489
2.313
1.922
30-Oct-12
1.796
1.620
2.104
1.968
31-Oct-12
1.663
1.541
1.976
1.877
01-Nov-12
1.683
1.581
1.985
1.908
02-Nov-12
1.714
1.631
2.005
1.939
03-Nov-12
1.673
1.621
1.966
1.908
04-Nov-12
1.704
1.671
2.005
1.942
05-Nov-12
1.785
1.750
2.113
2.067
Table 3 indicates that groundwater is hydrostatic up to 35 m depth and the piezometric level is at about 2 m below ground, which is consistent with other piezometer readings at the Jansen site.
3.3 Pile Capacity Estimation using Procedure (Canadian Foundation Engineering Manual 2006)
For cohesive soils with su>100 kPa, no clear methods for skin friction for bored pile were suggested by CFEM. Use of effective stress approach is probably more suitable. CFEM provided β and Nt values in the range of 0.25 to 0.32 and 3 to 10, respectively. Using an average β and Nt values of 0.28 and 7.0, Table 4 summarizes the estimated ultimate shaft and end bearing resistances with depth.
Table 4. Estimated ultimate shaft and end bearing resistances by CFEM suggested β-method
Depth (m)
1 σ’v (kPa)
fs (kPa)
fb (kPa) Nt=7
fb, (kPa) Nt=9
5
51
14
357
460
10
102
28.5
714
920
15
153
43
1,071
1,380
20
214
60
1,500
1,925
30
336
94
2,352
3,024
1 using average bulk unit weight of 20 kN/m3 up to 15 m depth and 22 kN/m3 below 15 m; groundwater table at ground surface.
An attempt is also made to calculate β using the relationship Ko tan δ, where Ko is the coefficient of earth pressure at rest (ie, 1-sinφ’). Using φ’ of 30 degrees for hard sandy silt till and δ = φ’, β becomes 0.29 which is very close to 0.28. The ultimate toe resistance is also calculated using Nt value of 9, as recommended by Tomlinson and Woodard (2009) and

6. Fleming et al (2008), which is also within the range of 3 to 10.
3.4 Pile Capacity Estimation using Procedure of FHWA (2010)
For cohesive soils, FHWA (2010) suggested the use of total stress approach (α-method) for shaft resistance calculation. Using the approximate su profile for the hard glacial till at the pile test site, Table 5 summarizes the ultimate shaft and end bearing resistances with depth following the FHWA (2010) recommendations.
Table 5. Estimated ultimate shaft and end bearing resistances by FHWA (2012) suggested α-method
Depth (m)
su, kPa
α
fs, kPa
fb, kPa (Nt=9)
5
175.5
0.53
92
1,580
10
135.5
0.55
75
1,220
15
245.0
0.46
112
2,210
20
355.0 1
0.35
124
3,200
1 due to limited data, su is assumed to be same below 20 m.
3.5 Estimation of Shaft Resistance using α-curve Developed by Weltman and Healey (1978)
Based on drilled shaft pile test data for piles installed in glacial till in the UK, Weltman and Healey (1979) proposed a correlation between α and su (Figure 3) which is significantly different than α-su correlations provided in both the CFEM and FHWA. Figure 3 summarizes the ultimate shaft resistances with depth using α values from Table 6.
Figure 3. Correlation between α and su, (Weltman and Healey, 1979)
Table 6. Estimated ultimate shaft resistances using α- curve (developed by Weltman and Healey 1978)
Depth (m)
su (kPa)
α
fs (kPa)
5
175
0.40
70
10
135
0.55
74
15
245
0.35
85
20
355
0.35
124
1 due to limited data, su is assumed to be same below 20 m.
3.6 Comparison of Predicted Values
3.6.1 Skin Resistance
Overall, values predicted by CFEM are low compared to the other two methods and values predicted by FHWA method are highest.
3.6.2 End Bearing
The predicted values using the methods suggested by CFEM and FHWA are comparable below 10 m. More comments on the different prediction methods and their suitability for prediction of pile capacity in hard glacial till at the subject site located in Saskatchewan will be provided below.
4 MEASURED PILE COMPRESSION CAPACITY
4.1 Summary of Compression Test Piles
Three CFA piles were tested in compression (C-1, C-2 and C-3) between March 15 to 17, 2012 by applying bi- directional load (Osterberg-Cell® (O-Cell®) testing). Table 7 summarizes the test pile information.
Table 7. Summary of test pile information (compression)
1 Strain gauge
2 The closer the SG to the base the higher the shaft resistance
4.2 Interpretation of Test Pile Data (compression)
A typical O-Cell® load vs displacement plot shows an upward and a downward movement vs load as the O- Cell® splits the pile shaft into two segments and applies load bi-directionally. The upward movement plot is the load-settlement curve of the pile shaft and the downward movement plot is the load-settlement curve of the pile base. Therefore, the ultimate skin resistance (fs) is mainly interpreted from the strain gauge (SG)
Test pile number and nominal diameter (mm)
Length and bottom elevation (m/masl)
Casing length and bottom elevation (m/masl)
O-Cell® depth and elevation (m/masl)
Number of 1SG above O-Cell®
Number of 1SGbelow O-Cell®
Max applied load (MN)
C-1/600
20.15/ 521.27
0.90/ 540.53
15.00/ 526.40
3 to +/- 3m spacing
Two (near the base)
2.24
C-2/750
25.10/ 516.38
0.80/ 540.65
17.50/ 523.98
4 to +/- 3m spacing
Two 2 (evenly distributed)
4.60
C-3/900
25.10/ 516.25
0.80/ 540.54
18.00/ 523.33
4 to +/- 3m spacing
Two (evenly distributed)
5.48

7. readings located above the O-Cell®. However, depending on the location of O-Cell®, fs is also interpreted from any SG readings located between O-Cell® and tip of pile for the portion of the pile below the O-Cell®. The ultimate base resistance (fb) is only interpreted using the SG located near pile base.
Some important observations from pile load tests are summarized below:
1. In general, fs increases with depth, even though the soil surrounding the pile is uniform, primarily hard till.
2. Above the O-Cell®, fs is the lowest within the softer till zones.
3. Higher shaft friction (nearly double) has been observed below 20 m in C-2/C-3 as compared to C-1; this could be due to the characteristics of the till changing below depths of 20 m to 22 m or due to the presence sand or gravel seams/layers
4. The back calculated α values range between 0.46 and 1.1; the α-su relationship is shown in Figure 4; this very closely mirrors the weltman-healy curve (virtually the same shape), except that this curve is shifted to the right (higher alpha value for lower su).
Figure 4. α-su relationship from measured data
4.3 Comparisons Between Estimated and Measured Pile Compression Resistance
Table 8 summarized the pile capacities predicted by the methods described above as compared to the measured capacities from the load tests.
Table 8. Summary of predicted vs measured pile capacities.
5 CONCLUSIONS
We offer the following conclusions/recommendations with respect to estimation of drilled shaft capacity in very stiff to hard glacial till in Western Canada:
1. All of the methods generally under estimated both skin resistance and end bearing. This could be due to the fact that existing methods (with the exception of the Weltman and Healey curve for shaft resistance) are generally based on pile load test data for purely cohesive soils. However, glacial till is a heterogeneous mixture of sand, silt, and clay, and is typically sandy or silty (sandy, silty till was prevalent at the pile load test site). This suggests that the very stiff to hard glacial till in Western Canada has a higher unit resistance for a given soil shear strength as compared to purely cohesive soils.
2. A single α-su curve, as found in most text books and foundation design guidelines, does not appear to be applicable for all cohesive soils and appears to under estimate the resistance of very stiff to hard glacial till soils in Western Canada.
3. The methods described in CFEM appear too significantly under estimate pile capacity for piles installed in the very stiff to hard glacial till in Western Canada.
4. The estimated skin resistances using the FHWA suggested α-method were generally closest to the measured values. As such, the FHWA method is recommended for design purpose for drilled piles installed in glacial till soils in Western Canada.
5. The estimated skin resistances based on the Weltman and Healey curve is the second closest to the measured values, and appears to be suitable for design purposes for drilled piles installed in glacial till soils in Western Canada. However, it is strongly suggested that efforts be made to develop a curve similar to that of Weltman and Healey for glacial till soils encountered in Western Canada.
6. It is possible that current geotechnical practice for evaluating pile design parameters for drilled shafts installed in western Canadian glacial till soils is too conservative.
Depth (m)
fs (kPa)
fb (kPa)
Measured
Estimated (% difference)
Estimated (Nt=9)
Measured
CFEM
FHWA
Weltman and Healey
CFEM
FHWA
5
(90+115+90)/3
= 98
14
(-85%)
92
(-6%)
70
(-27%)
460
1,580
-
10
(70+115+125) /3
= 103
28.5
(-72%)
75
(-27%)
74
(-28%)
920
1,220
-
15
(100+150)/2
= 125
43
(-82%)
112
(-13%)
85
(-40%)
1,380
2,210
-
20
(120+100)/2
= 110
60
(-45%)
124
(+ 12%)
124
(+ 12%)
1,925
3,200
3,250

8. ACKNOWLEDGMENTS
We would like to acknowledge Adam Lai, P.Eng. and Tara Stratton for their contributions.
REFERENCES
American Association of State Highway and Transportation Officials (AASHTO). 2007. “AASHTO LRFD Bridge Design Specifications, Customary U.S. Units, 4th Edition, Section 10, 'Foundations'”, Washington, D.C.
Bowles, J.E. 1996. Foundation Analysis and Design, 5th Edition. The McGraw-Hill Companies, Inc.
Canadian Foundation Engineering Manual (CFEM), 4th Edition. 2006. Canadian Geotechnical Society, co BiTech Publisher Ltd.
Federal Highway Administration (FHWA). 2010. Drilled Shaft Foundations (Publication No. FHWA-NHI-10- 016, May 2010).
Fleming, K; Weltman, A; Randolph, M. F.; and Elson, K. 2009. Piling Engineering, 3rd Edition. Taylor & Francis.
Kulhawy, F.H. and Chen, J.-R. 2007. “Discussion of ‘Drilled Shaft Side Resistance in Gravelly Soils’ by Kyle M. Rollins, Robert J. Clayton, Rodney C. Mikesell, and Bradford C. Blaise,” Journal of Geotechnical and Geoenvironmental Engineering, ASCE, Vol. 133, No. 10, pp. 1325-1328.
Prakash, S and Sharma, H.D. 1990. Pile Foundations in Engineering Practice. John Wiley & Sons, Inc.
Randolph, M. F. 2003. Science and Empiricism in Pile Foundation Design. Geotechnique, 53(10), 847-5- 875.
Skempton, A.W. 1951. The Bearing Capacity of Clays, in Building Research Congress. London: ICE, pp. 180–189.
Stas, C. V. and Kulhawy, F. H. 1984. Critical Evaluation of Design Methods for Foundations under Axial Uplift and Compression Loading. Report El-3771. Electric Power Research Institute, Palo Alto, California.
Terzaghi, K.; Peck, R.B., and Mesri, G. 1996. Soil Mechanics in Engineering Practice, Third Edition. John Wiley & Sons, Inc.
Tomlinson, M and Woodard, J. 2008. Pile Design and Construction Practice, 5th Edition. Taylor & Francis.
Weltman, A.J. and Healy, P.R. 1978. Piling in Boulder Clay and other Glacial Tills. DoE/CIRIA Report PG 5, London.
LIST OF SYMBOLS
α – coefficient relating unit side resistance to undrained shear strength, also called as adhesion factor)
B – pile diameter (m)
β – shaft friction coefficient
Df – depth of foundation (m)
δ – effective stress friction angle for pile-soil interface (degree)
φ’ – effective internal friction angle of soil (degree)
c – cohesion of soil (kPa)
fb – ultimate or unit toe resistance, also called end bearing (kPa)
fs – ultimate or unit pile shaft resistance, also called skin or side or shaft friction or resistance (kPa)
γ’ – effective unit weight of soil
K – coefficient of lateral earth pressure (= σ’h / σ’v)
Ko – coefficient of at rest earth pressure
Kp – coefficient of passive earth pressure
L – pile length (m)
N60 – SPT N value corrected for field procedures
Nt – bearing capacity factor, also denominated as Nc in some other papers / textbooks
Nc – bearing capacity factor, see Nt
Nγ – bearing capacity factor
Nq – bearing capacity factor (= γ’Df)
pa – atmospheric pressure (~101.43 kPa)QT – axial compression load applied to the butt (top) of the pile
su – undrained shear strength (kPa), also denominated as Cu in by Weltman and Healey (Figure 3)
σ’h – horizontal effective stress
σ’p – effective pre-consolidation pressure (kPa)
σ’v – vertical effective stress
z – depth (m)