DRILLED SHAFT CAPACITY IN COMPRESSION

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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.
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  • 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 Capa
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