SUGGESTION ABOUT DETERMINATION OF THE BEARING CAPACITY OF PILES ON THE BASIS OF CPT SOUNDING TESTS

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2 Figure 1. Measurement results of a Cone Penetration Test. Table 1. Correlation between skin friction and cone resistance. Pile type Timber Precast concrete Steel displacement Open ended steel tube + H-section Open ended steel tube driven into fine to medium sand Ultimate unit skin friction (τ s ) 0,012 q c 0,012 q c 0,012 q c 0,008 q c 0,0033 q c The average resistance against the progress of the cone, or penetration ( q c ), can be derived using the formula: qc 1 + qc 2 qc = (2) 2 In the Netherlands, in accordance with the advice of Meigh [3], the method generally applied is the one in which the average cone resistance ( q c 1 ) is determined to the depth of four times the pile diameter (4D) below the toe, and the average cone resistance ( q c 2 ) to the depth of eight times the pile diameter (8D) above the pile toe. Regarding the 4D 8D method, it is important to note that: minor peak depressions have to be ignored in the calculation; supposedly, they do not refer to thin weak strata, and the q c > 30 Mpa values will be also ignored in this interval. Obviously, there are also methods other than the 4D 8D method; in use, however, they only differ in the calculated depth below the pile toe (for example by taking 2D, instead of the 4D suggested above). Te Kamp [9] preferred to suggest the safety factors presented in Table 2, for the calculation of limiting capacity in the Netherlands, when the 4D 8D method is used: 46. ACTA GEOTECHNICA SLOVENICA, 2005/1

3 Table 2. Factor of safety for piles. Pile type Timber Precast concrete, straight shaft Precast concrete, enlarged shaft Factor of safety 1,7 2,0 2,5 Because of the disturbance and loosening of the soil via the boring tool, the Codes advise not to use cone resistance values when the skin resistance of bored piles is calculated. The relationship established for Dutch soil conditions is not necessarily applicable to cohesionless soils everywhere. The yielding and rupture of the soil caused by pushing a cone into the ground are different from those resulting from driving a pile by hammer followed by static loading. The work of Vesic [10] has shown the importance of the state of preconsolidation and mineralogy of the soil grains in any correlation of in-situ conditions with pile resistance. By coincidence, static cone resistance in the Netherlands (and Belgium) was found to be equal to pile base resistance. Elsewhere, Gregersen et al. [11] found the pile base resistance to be only one half of the cone resistance for loose medium to coarse sands in Norway, and Gruteman et al. [12] reported that a factor of 0.75 was applied to the cone resistance to obtain the ultimate base resistance of piles in silty sands in Russia. 2 IN-SITU TESTS The analysis of in-situ test data can result in better design parameter estimates which will affect the ultimate bearing capacity Q u of piles. A comparison of the Static Pile Load Test and CPT measurements in Hungarian soils was undertaken to better define mechanisms affecting Q u and to create formulae that are appropriate for Hungarian soils and that also consider construction methods. In this sense, the author selected the results of domestically performed Static Pile Load Tests where the results of the CPT tests were also available. Altogether, data from seven CFA tests, three tests with protective tubes, and 26 Franki piles were gathered (Table 3). 3 SUGGESTED FORMULAE TO CALCULATE THE ULTIMATE BEARING CAPACITY OF PILES The derived formula with the values in Fig. 2 relates to the failure load of a single pile. In deriving the formulae, the customary static basis has been used as a starting point, whereby the ultimate bearing capacity of a single pile (Q u ) is the sum of the ultimate resistance of the base (Q b ) and the ultimate resistance of the shaft (Q s ) capacities: Qu = Qs + Qb (3) Figure 2. Concept for the estimation of the failure load of a single pile. ACTA GEOTECHNICA SLOVENICA, 2005/1 47.

4 The first part of the formula (Q s ) depends on the total surface area of the shaft, the earth pressure acting thereon, the interactive forces between the surrounding soil and the shaft, and the technology of fabrication. These make: Q = β U L τ = β U A (4) s s s s fs q c = average value of the cone resistance below the pile toe (explained in Fig. 1) [MPa]. [It has been observed that the depth (n D) below the pile toe is strongly influenced by the applied piling technology, which has to be accounted for when the average q c value is derived]. where: U = length of the pile s periphery [m] L = length of the pile [m] A fs = area of the plotted f s curve from the CPT probe test (explained in Fig. 2) [MPa m] β s = empiric factor with view on the applied piling technology; it expresses the shaft resistance [-] τ s = average ultimate skin friction along the pile shaft [MPa]. The second part of the formula (Q b ) depends on the extension of the surface area where the pile toe rests, the specific resistance of the soil in the zone of the pile base, and on the applied piling technology. These make: where: Qb = b Ab qc α 60 β cos 2 (5) A b = nominal plan area of the base of the pile [m 2 ] β b = empiric factor with view on the applied piling technology; it expresses the base resistance [-] Table 3. Comparison of calculated and measured bearing capacities of piles. Location (ap = motorway) Pile type Length [m] Diameter [m] Soil below the toe 4 ANALYSIS OF THE ULTIMATE CAPACITY OF PILES USING THE STATIC PILE LOAD TEST The pile load tests were performed according to the standard loading procedure described in the Hungarian Standards, MSZ :1989 and MI :1984. All pile load tests have been carried out until the failure load was reached. 5 RESULTS OF THE COMPARISON OF THE MEASURED (STATIC PILE LOAD TEST) AND CALCULATED (CPT) ULTIMATE BEARING CAPACITIES OF PILES The results of the recommended CPT method used to estimate the ultimate bearing capacity of the selected piles discussed in the in-situ tests section were compared with the results of the Static pile Load Tests. The findings can be seen in Table 3. From CPT using the formulae Static Pile Load Tests Difference β s β b n Q u, calculated Q u, measured [1] [1] [1] [kn] [kn] [%] M3 ap./b 2 7,00 0,60 Gravel % M3 ap./b 3 5,00 0,60 Gravel % M3 ap./h 29 7,00 0,60 Gravel % 1,40 1,70 3 M30 ap./1 9,50 0,60 Gravel % M30 ap./4 7,00 0,60 Sand % Franki M3 ap./h 30 4,00 0,60 Sand % M3 ap./b 9 6,50 0,60 Clay % M3 ap./b 6 7,00 0,60 Clay % 2,40 2,70 3 M3 ap./b 7 9,00 0,60 Clay % M3 ap./b 11 9,00 0,60 Clay % M3 ap./b 13 23,00 1,00 Clay % Protective M3 ap./b 14 17,80 1,00 Clay 0,45 0, % tube M3 ap./h 32 20,60 1,00 Clay % M3 ap./hb 44 15,50 0,80 Sand % M3 ap./hb 46 14,50 0,80 Sand % M3 ap./hb 47 13,50 0,80 Sand % CFA 0,75 0,75 2 M3 ap./h 35 14,60 0,80 Sand % M3 ap./hb 42 15,80 0,80 Fine sand % M30 ap./2 13,80 0,80 Clay % 48. ACTA GEOTECHNICA SLOVENICA, 2005/1

5 Based on the results of the performed calculations, the regression coefficient (r) for each piling technology is as follows: For D = 60 cm diameter Franki piles: r = 0.87, For D = 100 cm diameter piles bored in protective tubes: r = 0.84, For D = 80 cm diameter piles bored with CFA technology: r = On the basis of piling technologies and pre-calculations the assumptions used and conclusions are as follows: In the course of the calculations the bulb diameter for the Franki piles was assumed to be equal to the trunk diameter; so the expansion of the bulb is included in the factor β b. To account for the densification of the soil in the case of driven D = 60 cm diameter Franki piles in granular soils, it is recommended to use the values β s = 1.40 and β b = 1.70 (higher than for the K o = 1-sinϕ equilibrium pressure), as well as 3D zone-depth, in the calculations. For D = 60 cm diameter Franki piles in cohesive soils, it is recommended to use β s = 2.40, β b = 2.70, and 3D zone-depth. The values β s = 2.40 and β b = 2.70 are just one unit higher (because of the pore-water pressure) than in the case of granular soils. In the case of D = 100 cm diameter piles bored in protective tubes presumably due to the accumulated pulverised sediment at the bottom of the hole it is recommended to use β b = 0.05 for the base resistance and β s = 0.45 for the shaft resistance (lower than K o = 1-sinϕ), as well as 1D depth-zone. For D = 80 cm diameter piles bored with the CFA technology, it is recommended to use β b = β s = 0.75 and 2D depth-zone. 6 SUMMARY This study presents the evaluation of a new method in predicting the ultimate bearing capacity of different piles (Franki piles, piles bored in protective tubes and piles bored with the CFA technology) driven into different soils in Hungary. Thirty six pile load test reports with CPT soundings adjacent to the test pile were collected. The prediction of pile capacity was performed for each pile; however, the statistical analysis and the evaluation of the suggested prediction method were based on the results of the nineteen piles (presented in this paper) that plunged (failed) during pile load tests. An evaluation scheme was executed to evaluate the CPT values based on the ability to predict the measured ultimate bearing capacity. Different values (β s, β b and n D) were suggested for different piling technologies for the evaluation scheme. Based on the results of this study, the suggested formulae using the results of the CPT testing are given to predict the ultimate load bearing capacity of the piles. While one may not expect that any calculation carried out based on the result of either the CPT test or any other probing test will lead in all cases straight to the determination of the exact ultimate bearing capacity of piles derived by using static loading test results, the performed study proves that more accurate approaches can be found to replace traditional static formulae in the design stage. REFERENCES [1] Chen, B.S.-Y. and Mayne (1995). Type 1 and 2 piezocone evaluations of overconsolidation ratio in clays. Proceedings, International Symposium on Cone Penetration Testing (CPT 95), Vol. 2., Swedish Geotechnical Society Report No. 3:95, Linkoping, [2] De Beer, E. (1963). Scale effect in the transposition of the results of deep sounding tests on ultimate bearing capacity of piles and caisson foundations. Geotechnique 23, 39. [3] Meigh, A. C. (1987). Cone penetration testing, CIRIA-Butterworth. [4] Meyerhof, G. G. (1976). Bearing capacity and settlement of pile foundations. Journ. Geot. Eng. Div., ASCE (102) GT 3, [5] Poulos, H. G. (1989). Pile behavior-theory and application. Geotechnique, Vol. 39, No. 3, [6] Titi, H. H. (1999). Evaluation of bearing capacity of piles from cone penetration test data. Louisiana Transportation Research Center, LA. [7] Tomlinson, M. J. (2001). Foundation Design and Construction. 7th ed., Pearson Education Ltd, Essex, [8] Tomlinson, M. J. (1971). Some effects of pile driving on skin friction, Proceedings of the Conference on the Behavior of Piles, Institution of Civil Engineers, London, ACTA GEOTECHNICA SLOVENICA, 2005/1 49.

6 [9] Te Kamp, W. C. (1977). Sondern end funderingen op palen in zand. Fugro Sounding Symposium, Utrecht. [10] Vesic, A. S. (1977). Design of pile foundations. NCHRP Synthesis 42, Transportation Research Board, Washington D. C. [11] Gregersen, O. S., AAS, G. and Dibiagio, E. (1973). Load tests on friction piles in loose sand. Proceedings of the 8th International Conference, ISSMFE, Moscow, Vol. 2.1., [12] Gruteman, M. S. et al. (1973). Determination of pile resistance by means of large-scale probes and pile foundation analysis based on allowable settlements. Proceedings of the 8th International Conference, ISSMFE, Moscow, Vol. 2.1., ACTA GEOTECHNICA SLOVENICA, 2005/1

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