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FOUNDATION DESIGN WITH MÉNARD PRESSUREMETER TESTS
French contributions to
International
Foundation Congress & Equipment Expo '09
Excerpts from Contemporary Topics in In Situ Testing, Analysis, and Reliability of Foundations, Proc. Int. Foundation Congress and Equipment Expo ’09 (IFCEE’09), Orlando, Florida, 15-19 March 2009, ASCE Geotechnical Special Publication No. 186.
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SUMMARY DESIGN RULES 1.
Direct Design Rules For Piles Using Ménard Pressuremeter test Michel (Mike) Gambin, Fellow ASCE & Roger Frank
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ASCE Geotechnical Special Publication No. 186 p.111-118
2. Pile Design at Failure Using the Ménard Pressuremeter : an Up-Date
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Michel Bustamante, Michel (Mike) Gambin, Fellow ASCE & Luigi Gianeselli ASCE Geotechnical Special Publication No. 186 p.127-134
3. ISP5 Pile Prediction Revisited
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Philippe Reiffsteck, A.M. ASCE ASCE Geotechnical Special Publication No. 186 p.50-57
CASE HISTORIES 4. Rades Bridge Drilled Shafts Designed and Tested Using
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Ménard Pressuremeter François Schlosser, Alain Guilloux, Kamel Zaghouani, Patrick Berthelot ASCE Geotechnical Special Publication No. 186 p.42-49
5. LNG Tanks at Damietta on Drilled Shafts Designed and Tested Using Ménard PMT Jean-Yves Boumedi, Jean-Pierre Baud & Bruno Radiguet ASCE Geotechnical Special Publication No. 186 p.103-110
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Direct Design Rules For Piles Using Ménard Pressuremeter Test Michel (Mike) Gambin1, Fellow ASCE & Roger Frank2 1
Scientific Adviser , Apageo, 21 quai d’Anjou, 75004 Paris, France,
[email protected] Professor, Université Paris-Est, Ecole nationale des ponts et chaussées, Navier-CERMES, Cité Descartes, Champs sur Marne, France 2
ABSTRACT Direct design rules derived from PMT data are used for estimating the bearing capacity and settlement of piles and for their behaviour under lateral loading. The theoretical background of these rules is explained by comparing borehole expansion by the pressuremeter to soil response under the various pile displacements. The most recent developments are submitted. 1. INTRODUCTION The Ménard pressuremeter is probably the instrument which most closely models the way soil behaves around a pile (Baguelin et al. 1978). It yields a failure parameter and a small strain (10-2) deformation parameter. It is a particularly good tool to analyze axial bearing capacity, pile settlement and behaviour under lateral loading (Briaud 1995). 2. THE MÉNARD PRESSUREMETER A Ménard pressuremeter test (PMT) is somewhat different from other in-situ tests such as SPT or CPT which are or were originally used in geotechnical design on the basis of correlations. In a MPM test a cylindrical cavity, typically at one meter intervals of depth, in any type of soil from soft soil to weak rock, is subjected to pressure increments and the resulting expansion is measured in terms of a volume increase. Louis Ménard’s question was to ask: why take a sample from a borehole, bring it to a lab and test it, sometimes in poor condition and certainly having undergone a stress reversal, when it is possible to insert an instrument into a predrilled hole and to carry out a plane strain loading test on the in-situ soil? Since the hole is very small (60mm), caving is unlikely and, if it could occur, drilling with a mud as in the oil industry is an excellent way to create a suitable cavity. Since it is impossible in a simple test to stress the soil vertically as the structure will do testing is by applying a known lateral pressure on the vertical wall of the
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borehole. This, however, stresses the soil in three dimensions anyway and at the actual in-situ stress. A test consists of 4 pairs of pressure and volume readings at each pressure holds at 1, 15, 30 and 60 seconds. The number of pressure holds for one test is always greater than 7 and number of readings during a test is thus more than 50. Data loggers print out values of the main parameters and graphs on the spot. Since tests are almost always carried out at one meter depth intervals a lot of data is obtained from a single borehole. 3. THEORY BEHIND THE MÉNARD DESIGN RULES By a simple analysis of the continuous stress-strain curve obtained at each test depth, two main parameters characterize each soil layer (ASTM 1987 - present) : - an E-modulus called Pressuremeter Modulus, written EM, and - a Limit Pressure, written pLM , which by convention is reached when the volume of the expanded cavity has doubled. The determination of pLM requires a simple graphical operation, EM, however must be derived from the equation for the expansion of an infinite cylindrical cavity in an elastic medium which produces the shear modulus G from the shear deformation of the pressuremeter test:
ΔR / R = [ 1 / 2G ] Δp where ΔR / R is the radial strain and Δp is the corresponding increase in the applied pressure From which E can be derived using: EM = 2 ( 1 + Q G
(2)
Poisson’s ratio Qis conventionally taken equal to 0.33 Note that the parameters c’ and I’ cannot easily be obtained since, unlike a confined tri-axial soil sample, each concentric ring of soil around the pressuremeter is at a different level of stress and strain. Bearing Capacity of Piles and Drilled Shafts The theory, due to Prandtl, using c’ and I’ to assess the bearing capacity of a shallow footing is recognised as only approximate, since, in terms of the Rankine passive earth pressure theory, the soil is assumed to behave as a rigid-plastic body. The soil response in a pressuremeter test behaves much closer to the way soil reacts to a loaded deep foundation a fact observed by many researchers e.g. Skempton et al., (1953), Vesic, (1972) & (1977), Salgado & al. (1997), and, of course, Louis Ménard (1963) who not only originated the theory, but checked it against full scale tests on prototype footings and piers.
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A simple verification is this (Fig.1) : (a) to estimate the soil bearing failure under shallow strip footings or spread footings, analysis using c’ and I’ gives good results but, since the ultimate bearing capacity is almost always a direct function of the size of the footing, it is the estimate of the settlement, which provides the value of the bearing stress in service (Terzaghi & Peck, 1948), especially in sands, (b) when deeper foundations are considered, as when piers and piles have to be designed, conventional theory does not permit modelling the actual soil failure below a pile tip: The soil around the tip is in a plastic state and the soil reaction around this sheared volume can be compared with the response to the expansion of a deeply embedded cavity. Elastic/plastic theory must be used, where the elastic response
FIG.1 Shallow foundation bearing failure against pile tip bearing failure in an homogeneous soil (EM & pLM constant). of the soil outside the sheared volume dominates. Ménard (1963) shows that there is a simple theoretical relationship between the soil failure stress below the tip of a pile qL and the limit pressure pLM measured at completion of a pressuremeter test. Beginning from the work of Bishop et al, (1945), he could write : qL - qo = kp (pLM – po)
(3)
where qL is the ultimate bearing stress at the pile tip qo the vertical overburden stress at pile tip depth kp the Ménard Bearing Factor for this pile tip in this type of soil pLM the Ménard limit pressure at pile tip depth po the insitu horizontal effective stress at pile tip depth and it appears that, below a “critical depth”, the tip bearing capacity alone is much less than predicted by the c’ and I’ theory (Fig.1). Similarly, he also showed that a relationship between q s, the maximum skin friction resistance at a given depth, and pLM at the same depth can be written in the form of
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qs = f ( pLM, soil type & pile/shaft type).
(4)
These are correlations which rapidly started to receive confirmation based on hundreds of observations carried out by the French Government Laboratories for Bridges and Roads (LPC’s) on all types of prototype piles and drilled shafts using strain gauges along the whole pile shaft (Bustamante et al. 1981 - 2009). From this the ultimate total axial pile capacity Q can be expressed by Q = A kp (pLM – po) + P Σ (qsj . zj )
(5)
where A is the base area of the pile P the perimeter of the pile cross section zj the thickness of the jth soil layer exhibiting a uniform qsj skin friction. The EUROCODE 7-2, in its Annex E, § E.3 “Example of a method to calculate the compressive resistance of a single pile” (CEN, 2006) also quotes the LPC’s method. Pile Settlement Prediction For soil settlement below shallow footings, Ménard showed that this settlement is mostly governed by the deviatoric component of the stresses and not by the isotropic component. He derived an equation w = f (EM) involving both shear and, to a smaller extent, compression deformation, but also which took into account the degradation of the shear modulus G with increasing strain (Ménard 1961, 1962 & 1965). This approach could be extended to the settlement of piles using the method now called the load transfer method (Gambin 1963, Frank & Zhao 1982), in which the pile is divided into a series of equal length elements as in Fig.2. For the pile tip, the shear stress effect controls the settlement even more than below a footing since the expansion of the spherical cavity ΔR/R around the plastic body of soil is similar to the pressuremeter equation (1) :
ΔR/R = [1/4G] Δp
(6)
Thus, the settlement zp of a pile tip is simply given (Frank & Zhao 1982) by : zp = ( B/λp ) qp
(7)
where B is the diameter of the pile qp the pile tip pressure, with qp < qL, (Fig.2) and λp is a factor which varies from 4.8 EM in coarse soils up to 11 EM in fine soils up to a stress of qL /2 and is 5 times smaller above the qL / 2 stress. The skin friction qsi mobilized during the settlement of the ith pile shaft element is then obtained as a function of zsi the local shear displacement of this shaft element
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against the adjacent soil layer. This function involves EM (Frank & Zhao 1982) as
FIG. 2 Load transfer method to estimate pile settlement (Frank & Zhao, 1982). follows : zsi = ( B/λs ) qsi
(8)
where B is the diameter of the pile, qsi the estimated shaft friction of the ith element against the soil but limited by to a maximum of qs as shown in Fig.2 – values are given in Bustamante et al. (2009) and λs a factor which - varies from 0.8 EM in coarse soils to 2 EM in fine soils up to a stress of qs /2 - and is 5 times smaller above the qs / 2 stress. Results of comparisons between prediction and observation of pile load tests were given in a number of previous papers. In these Proceedings Boumedi et al. (2009) and Schlosser et al.( 2009) confirm them as satisfactory. Pile Lateral Loading Behaviour There are various cases where piles are subjected to lateral loading : - either the pile may be subject to a force or a moment or both at its head , as in a mooring dolphin or
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- the pile can be subject to the thrust of a soil layer displaced under an embankment, as an abutment pile. In both cases, the ground stress is similar to that observed during a pressuremeter test. Since the Winkler theory for horizontal beams on elastic supports EI d4y / dz4 + k.B y = 0
(9)
can be used, the value of k is readily obtained from the settlement equation w = f(EM) given by Ménard (1962) for a infinitely long strip footing, B in width : k.B = p/w. When EM values are averaged, for B larger than 0.6 m (2 ft), and below the critical depth : k.B = Es = EM {18 / [4 (2.65 B/Bo)α Bo/B + 3α]}
(10)
where α is the Ménard rheological factor ( 1/4 < α < 2/3) and Bo a reference diameter equal to 0.6 m.
FIG. 3 Soil reaction against lateral displacement for actions at head level (after Frank 1999) a) permanent forces at pile head ; b) soil lateral thrust ; b) short time forces at pile head; d) unexpected instant forces at pile head. This was checked using pressuremeter test data on various laterally loaded prototype piles (Gambin 1979). The present design rules (Frank 1999) using generalized P-y curves include the degradation of k when y increases (Fig.3). Finally, when a soil applies a horizontal thrust on the pile (Fig.4), the last term of the Winkler equation (9) must be replaced by
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EI d4y / dz4 + k.B [y(z) – g(z)] = 0
(11)
. FIG. 4 A pile subject to earth pressure (after Frank 1999) where y(z) is replaced by [y(z) – g(z)], g(z) being the free horizontal displacement of the soil in the absence of the pile. It is assumed that at equilibrium, for given values of applied forces and moments at pile head and pile tip, g(z) plays a role similar to that of y(z) in equation (8). Baguelin et al. (1978) have shown that the application of these design methods is satisfactory. CONCLUSION The present paper has tried to show that Ménard direct design rules are not a “black box”. The rules are based on a novel but rigorous approach to Soil Mechanics. In the search for sustainable developments and savings in materials, structural deformations cannot be overlooked as they used to be. Stiffness of soil has become as important as its strength. Since pressuremeter testing delivers 2 parameters associated with strength and stiffness the Ménard pressuremeter appears the ideal tool for our profession in this new century (Baker 2005). REFERENCES ASTM D-4719 (since 1987). “Standard Test Method for Prebored Pressuremeter Testing in Soils.” 2008 Annual Book of ASTM Standards, Vol. 04-06. Baguelin, F., Jézéquel, J.-F., Shields, D.H. (1978). The Pressuremeter and Foundation Engineering, Trans Tech Publications, Clausthal, 617 pages.
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Baker, C. N. (2005) “The Use of the Ménard Pressuremeter in Innovative Foundation Design from Chicago to Kuala Lumpur.” ISP5, Proc. Int. Symp., LCPC, Paris Bishop, R.F., Hill, R., Mott, N.F. (1945). The theory of indentation and hardness test, Proc. Physical Society, No.57, London. Briaud, J-L., (1992). The Pressuremeter, A. A. Balkema, Brookfield VT. Bustamante, M. & Gianeselli, L. (1981). “Observed and Predicted Bearing Capacity of Isolated Piles Using the Pressuremeter Method.” (in French) Revue Française de Géotechnique, No.16 Paris Bustamante, M., Gambin, M., Gianeselli, L. (2009). “Pile Design at Failure Using the Ménard Pressuremeter : an Up-Date.” Proc. IFCEE ’09, ASCE. CEN (2006). Eurocode 7 Geotechnical design – Part 2: Ground investigation and testing, pp. 123-124, Brussels. Clarke, B.G. (1995). Pressuremeters in Design, Blackie Academic and Professional [now :Taylor & Francis], London. Frank, R. (1974). Theoretical Study of Axially Loaded Piles - Introducing Dilation. (in French) Rapport de Recherche No.46, LCPC Paris Frank, R (1999). Design of Shallow and Deep Foundations (in French), Presses des Ponts, Paris, p. 80 – 123. Frank, R & Zhao, S. R. (1982). “Estimating the Settlement of Axially Loaded Bored Piles in Fine Sand by PMT Data.” (in French) Bull. Liaison LPC No.119 Paris Gambin, M. (1963). “Estimation of the Settlement of a Deep Foundation in Terms of Pressuremeter Tests Data.” (in French), Sols-Soils No.7, 1963 Paris Gambin, M. (1979). “Calculation of Foundations Subjected to Horizontal Forces Using Pressuremeter Data.” Sols-Soils No.30-31, Paris Ménard, L. (1961). “Influence of Stress Level and Stress History on Settlements.” (in French) Proc. Vth ICSMFE, 1/42 pp.249-253, Dunod Publisher, Paris Ménard L. (1962). “Evaluation of Settlements.” (in French) Sols-Soils No.1. Ménard, L. (1963). “Calculation of the Bearing Capacity of Foundations Based on the Results of Pressuremeter Tests.” (in French) Sols-Soils No.5 & 6. Ménard L. (1965). “Rules for Estimating Bearing Capacity and Settlement of Foundations Using PMT Data.” (in French) Proc. VIth ICSMFE, 4/15 pp.295-299. Salgado, R. Mitchell, J. & Jamiolkowski, M. (1997). “Cavity Expansion and Penetration Resistance in Sand.” J. Geotech. Eng. Vol.123, No.4 Skempton, G. W. Yassin, A. A. & Gibson, R. E. (1953). “A Theory for the Bearing Capacity of Piles in Sand.” (in French), Proc. (2nd) ECSMFE, Annales de l’ITBTP No.63-64, Paris Terzaghi K. & Peck R. (1948). Soil Mechanics in Engineering Practice, Wiley, NewYork (art. 54) Vesic A. S. (1972). “Expansion of Cavities in Infinite Soil Mass.” J. SM&F Div. Vol. 98 No. SM3, pp.265 - 290 Vesic, A. S. (1977). “Design of Piles.” U.S. Transp. Res. Board, NCRP Synth. No.42 Washington D.C.
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Pile Design at Failure Using the Ménard Pressuremeter : an Up-Date Michel Bustamante1, Michel (Mike) Gambin2, Fellow ASCE & Luigi Gianeselli3 1
CEO, MB Fondations, Saint Cloud, France,
[email protected] Scientific Adviser, Apageo, 21 quai d’Anjou, 75004 Paris, France,
[email protected] 3 Senior Engineer : MB Fondations, Saint Cloud, France,
[email protected] 2
ABSTRACT : The paper summarises the results of 30 years of pile loading tests on prototype piles installed by more than 26 different techniques and in which the soil was previously characterised using the Ménard pressuremeter. The present paper is based on the analysis of 561 load tests on more than 400 piles instrumented to record the limit unit skin friction of each separate soil layer and the limit end bearing. These are then compared with the PMT direct design rules initiated by Louis Ménard in the 1960’s. These rules are based both on the theory of cavity expansion in soils and on his own experiments. In a companion paper to this conference this method is applied to pile settlement prediction and to the design of piles subjected to lateral loading. 1. INTRODUCTION Since the early 1990’s, when the new French Code of Practice for Foundations (M.E.L.T. 1993), known as “Fascicule 62-V”, was published (Bustamante & Frank 1999), additional experimental data have been gathered by the LCPC, the French Highways Agency. These data include comprehensive site investigations with PMT, CPT and SPT. The plan was to test instrumented piles up to 2 m in diameter and : 1) to include the most recent installation techniques which are now common practice, 2) to refine the values in the analysis of the limit unit skin friction qs and the pile tip bearing factor kp. The aim was greater simplification whilst preserving the essentials of the method. 2. THE DIRECT DESIGN MÉNARD PMT METHOD The principles of this method are given in another paper to this Conference (Gambin & Frank 2009). During most pile load tests, the end of the test occurs when the pile head begins rapid subsidence. The load at this threshold is called the limit load QL. QL is defined as being the load at which the head settlement sL is given by sL t B/10 + 'e, where B is the diameter of the pile, and 'e is the pile elastic shortening.
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Since the bearing capacity of a pile is expressed by Q = A kp [(pLM – po)e] + P Σ (qsi . zi )
(1)
where A is the pile tip area kp the tip bearing factor [(pLM – po)e] the net equivalent Ménard limit pressure under the pile tip P the perimeter of the pile cross section the thickness of the soil layer ‘I’ exhibiting a uniform skin zi friction, qsi parameters kp and qs, which are essential to this equation, are measured on prototype piles instrumented with removable extensometers (Fig.1) and load tested to failure (Bustamante & Gianeselli 1981).
FIG. 1. Assembling a removable extensometer before insertion into a bored pile. By recording the strain gauge readings during the pile load test, it is possible to obtain the values of both qs for each soil layer along the shaft and kp for the pile tip (Bustamante & Doix, 1985), as shown by Reiffsteck (2009) in his Fig.4. 3. THE EXPERIMENTAL SUPPORT FOR THE UP-DATED RULES The Geotechnical Calibration In the previous papers the data were obtained from a total of 204 sites at which site investigations involved PMT’s, CPT’s, sometimes SPT’s and also lab tests on cored samples. It is interesting to analyze the chance of success of the various investigation techniques at depth to provide the required data (pLM, qc, N, or c’ and I’). The main soil categories investigated are clay, silt, sand, gravel, chalk, marl, marly limestone and weathered or fragmented rock. Although pile load tests were also carried out in other types of soil such as coral, volcanic and collapsible soils, swelling soils, etc., results are not yet sufficiently complete to derive specific design rules for them.
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Table 1 shows that for a large number of soil types in which piles are embedded (weathered or fragmented rocks, hardened or very fine cohesionless formations), the Ménard pressuremeter remains the most versatile site investigation tool. Table 1. Feasibility of in situ tests or coring at 204 sites.
Type of test
Number of sites as a function of test feasibility 1 Tests Tests Tests Insufficient Possible but 2 3 Completed No. of Tests Inadequate 4 Curtailed 155 Sites 3 Sites 46 Sites 0 Sites (76%) (1.5%) (22.5%) (0%) 60 Sites 79 Sites 23 Sites 42 Sites (29.4%) (38.7%) (11.3%) (20.6%) 26 Sites 54 Sites 72 Sites 52 Sites (12.7%) (26.5%) (35.3%) (25.5%)
PMT (pLM) CPT (qc) SPT (N) Coring for 21 Sites 67 Sites 69 Sites 47 Sites Laboratory (10.3%) (32.8%) (33.8%) (23.1%) (c’ and I’) 1 It is assumed that a PMT or an SPT log includes a test every meter. 2 Throughout the whole pile depth at least. 3 Insufficient No. of tests (PMT), premature refusal (CPT), excessive blow count (SPT) or sample badly recovered. 4 Tests deemed inadequate beforehand due either to soil type or to soil resistance. The Various Piles Analyzed Our up-dated analysis identified 26 basic pile installation techniques as opposed to 17 types for the French “Fascicule 62-V” (MELT 1993). These techniques are set out in Table 2. Techniques with common factors are now grouped under the same code number. This helps to choose the tip bearing factor k p. Among the 408 pile and anchor loading tests recently analyzed, 180 tests (or 44%) are related to piles which do not appear in Fascicule 62-V. They were described in five important papers (Bustamante & Gianeselli 1993 and 2005; Bustamante et al., 1991, 1998 and 2002). Out of a total of 561 tests to date, 276 tests (or 49%) could be taken to the limit load. For the remainder, the load was extrapolated up to this value by one of the usual analytical methods (Borel et al. 2004). Finally, 13% of the piles were subjected to tensile tests. 4. CHOOSING kp AND qs The use of the tables for kp and qs, need some explanations. The Tip Bearing Factor kp Parameter kp value can be chosen from Table 3 once the pile group code number is known. Since we now have more pile types in Table 2 we can select a single value of
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kp per pile type in Table 3. Furthermore there is no need now to apply a reducing factor for steel piles (Pile codes Nos. 5-7). Table 2 – Description and Characteristics of 418 Analyzed Piles. Group Type Pile2 Code No. Qty 1
8
B3 (mm) 5002,000 2701,800 2701,200 4201,100
D4 (m) 11.5-23
Pile Description Pile or Barrette Bored in the dry
Pile and Barrette Bored with Slurry 1 Bored and Cased Pile 3 2 20-56 (permanent casing) Bored and Cased Pile 4 28 5.5-29 (recoverable casing) Dry Bored Piles / or Slurry 51 4 520-880 19-27 Bored Piles with Grooved Sockets / or Piers (3 Types) Bored Pile with a single or a 1 2 6 50 410-980 4.5-30 double-rotation CFA (2 types) 7 38 310-710 5-19.5 Screwed Cast-in-Place 3 8 1 650 13.5 Screwed Pile with Casing Pre-cast or Pre-stressed 6.530 280-520 91 Concrete Driven Pile (2 types) 72.5 Coated Driven Pile 4 10 15 250-600 8.9-20 (concrete, mortar, grout) 11 19 330-610 4-29.5 Driven Cast-in-Place Pile 12 27 170-810 4.5-45 Driven Steel Pile, Closed End 5 13 27 190-1,22 8-70 Driven Steel Pile, Open End 14 23 260-600 6-64 Driven H Pile 6 15 4 260-430 9-15.5 Driven Grouted 5 or 6 H Pile 7 16 15 3.5-2.5 Driven Sheet Pile 17 2 80-140 4-12 Micropile Type I 1 18 8 120-810 8.5-37 Micropile Type II SGP 5 Micropile (Type III) / 10019 23 8.5-67 or SGP Pile 1,220 8 MRP 6 Micropile (Type IV) / 20 20 130-660 7-39 or MRP Pile 1 Some types may include several sub-types 2 Some piles subjected to several tests. 3 Minimum and maximum nominal diameter B. 4 Minimum and maximum full embedment depth D. 5 involving a Single Global Post grouting. 6 with Multiple Repeatable Post grouting. 2
64
6-78
The Ultimate Unit Skin Friction qs
-
Parameter qs is from Tables 2 and 4: (i) select the pile type from Table 2 and (ii) find the Qi applicable as a function
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of soil type in Table 4, Table 3. Values for the Tip Bearing Factor kp Group Clay Sand, Marl and Chalk Limestone Code & Silt Gravel 1 1.25 1.2 1.6 1.6 * 2 1.3 1.65 2.0 2.0 3 1.7 3.9 2.6 2.3 4 1.4 3.1 2.4 2.4 * 5 1.1 2.0 1.1 1.1 * 6 1.4 3.1 2.4 1.4 * 7 1.1 1.1 1.1 1.1 * 8 1.4 1.6 1.8 1.8 * A higher kp value can be used but must be proven by a load test
Weathered Rock 1.6 2.0 2.3 2.4 * 1.1 * 1.4 * 1.1 * 1.5*
Table 4. Selecting the Qi line to obtain the limit unit skin friction values qs Marl,
Weathered Rock Limestone 1 Q2 Q2* Q5 Q4 Q6** 2 Q2 Q2 Q5 Q4 Q6** 3 Q1 Q1 Q1 Q2 Q1** 4 Q1 Q2 Q4 Q4 Q4** 5 Q3 Q3* Q5 Q4 Q6 6 Q2 Q4 Q3 Q5 Q5** 7 Q3 Q5 Q4 Q4 Q4** 8 Q1 Q2 Q2 Q2 Q2** 9 Q3 Q3** Q2 Q2** (a) 10 Q6 Q8 Q7 Q7 (a) 11 Q2 Q3 Q6** Q5** (a) 12 Q2 Q2** Q1 Q2** (a) 13*** Q2 Q1 Q1 Q2 (a) 14*** Q2 Q2 Q1 Q2** (a) 15*** Q6 Q8 Q7 Q7 (a) 16*** Q2 Q2 Q1 Q2** (a) 17 Q1 Q1 Q1 Q2 Q6** 18 Q1 Q1 Q1 Q2 Q6** 19 Q6 Q8 Q7 Q7 Q9** 20 Q9 Q9 Q9 Q9 Q10** * If ground properties permit. ** Use of a higher value must be proven by a load test. *** Cross section and perimeter estimated according to Fig.3. (a) For pile groups No.9 – 16 and if rock condition permits penetration, choose the qs value proposed for marl and limestone or a higher one if this can be proven either by a load test or by reference to an existing example in the same local area. Pile Type No.
Clay, Loam
Sand, Gravel
Chalk
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-
(iii) use Figure 2 to obtain on the selected Qi curve the qs for the Ménard limit pressure pLM measured at the same depth.
The graph in Fascicule 62-V for the upper lines (then Q6 – Q7) shows a set of discontinuous straight lines. In Fig.2, the same qs lines (now Q6 – Q10) are continuous, which avoids any ambiguity when choosing this parameter. 0,70
qs
0,65 0,60
Q10
(MPa) Q9
0,55
Q8
0,50 0,45
Q7
0,40
Q6
0,35 0,30 0,25
Q5
0,20
Q4
0,15
Q3
0,10
Q2
0,05
pLM (MPa)
Q1
0,00 0,0
0,5
1,0
1,5
2,0
2,5
3,0
3,5
4,0
4,5
5,0
5,5
6,0
FIG. 2. Direct Design using PMT Data. Chart for unit skin friction qs Additional recommendations Most of the recommendations given in the current Code of Practice (MELT 1993) are valid for the use of Tables 2, 3 and 4 and the Chart in Figure 2. For driven piles areas and perimeters must be calculated according to Figure 3. For vibrated piles kP and qs must be reduced by a factor of 0.5 and 0.3 respectively (Borel et al. 2002). For more information about grouted piles and micropiles, the reader can consult the paper by Bustamante & Doix (1985). Finally, to design piles in hard soils PMT data should be obtained from high pressure equipment (Massonnet 2005).
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FIG. 3. Areas A and Perimeters P to be used for open-end steel piles & sheet piles 5. VALIDITY OF THE UP-DATE All the previous factors were checked by using them in reverse to calculate the ratio ‘QL measured / QL calculated’. Some results are given in Table 5. Table 5. Measured QL / Calculated QL Ratios. All Types of Piles 1 No. of piles 204 Mean 1.020 Standard error 0.009 Median 1.018 Standard deviation 0.124 Variance 0.015 Screwed Cast in Place Piles No. of piles 38 Mean 1.029 Standard error 0.016 Median 1.026 Standard deviation 0.099 Variance 0.009 1
Bored Piles 2 No. of piles 37 Mean 1.047 Standard error 0.020 Median 1.052 Standard deviation 0.119 Variance 0.014 Grouted Piles and Micropiles3 No. of piles 19 Mean 0.980 Standard error 0.029 Median 1.011 Standard deviation 0.127 Variance 0.016
Only loading tests strictly carried out to QL and excluding tension tests Sub-set of the previous group 3 Involving either a Single Global Post grouting or Multiple Repeatable Post grouting. 2
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6. CONCLUSION Up-dating the Direct Design Ménard Pressuremeter Method for calculating the limit pile load QL led to : 1) an adjustment of the parameters qs and kp for a total of 26 different pile types 2) a simplification in the number of tip bearing factors kp for each soil category and pile technique; 3) the proposal of a chart involving 10 continuous qs curves incorporating the most recent pile techniques. 8. REFERENCES Borel, S., Bustamante, M., Gianeselli, L. (2004). “An appraisal of the Chin method based on 50 instrumented pile tests”, Ground Engineering, January, Vol.37, No.1, pp.22-26. Bustamante, M., Borel, S., Gianeselli, L., (2002). “Two comparative field studies of the bearing capacity of vibratory and impact driven sheet piles”, Proc. TRANSVIB, 19-21 March, Louvain-la-Neuve, Belgium, Balkema. Bustamante, M., Doix, B. (1985). “Design Method for Ground Anchors and Grouted Micropiles” (In French) Bull. Liaison Labo. P. et Ch. No.140, pp.75-92. Bustamante, M., Frank, R. (1999) “Current French Design Practice for Axially Loaded Piles” Ground Engineering March, London, pp.38 – 44. Bustamante, M. & Gianeselli, L. (1981) “Observed and Predicted Bearing Capacity of Isolated Piles Using the Pressuremeter Method” (in French) Revue Française de Géotechnique, No.16, Presses des Ponts, Paris Bustamante, M., Gianeselli, L., (1993). “Design of auger displacement piles from in situ tests", 2nd Intern. Geotech. Seminar: Deep Foundations on Bored and Auger Piles, Balkema. Bustamante, M., Gianeselli, L. (2005). “Design of Screwed Piles with Ménard Pressuremeter” (in French). Proc. ISP5, 22-24 August, Presses des Ponts, Paris, Vol. 1, pp.447-456. Bustamante M., Gianeselli L., Koch G. (1991). “Vertical Bearing Capacity of SheetPiles” (in French), Proc. Col. Inter. Fondations Profondes, Presses des Ponts, Paris, pp.145-152. Bustamante, M., Gianeselli, L., Weber, L., (1998).“The bearing capacity of driven steel piles in weathered chalks”, Proc. 7th Int. Conf. and Ex. on Piling and Deep Foundations, DFI 98. Gambin, M. (1963). “The Ménard Pressuremeter and the Design of Foundations”(in French) Actes Journées des Fondations, Laboratoire Central des Ponts et Chaussées, Paris. Gambin, M. and Frank, R. (2009), "Direct Design Rules for Piles Using Ménard Pressuremeter, Proc. IFCEE ‘09, ASCE. Massonnet, R., (2005). “High Pressure Ménard Pressuremeter” Proc. ISP5, Presses des Ponts Paris, pp. 81-90. M.E.L.T. (1993) Design Rules for Foundations, Tender Documents for Public Works, Fasc. No.62, Titre V (in French), Imprimerie Nationale Paris.182 pages. Reiffsteck, P. (2009) ISP5 Pile prediction revisited, Proc. IFCEE 09.
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ISP5 Pile Prediction Revisited Philippe Reiffsteck1, A.M. ASCE 1
Researcher, Université Paris Est, Laboratoire Central des Ponts et Chaussées, 58 Bd Lefebvre, 75015 Paris, France;
[email protected]
ABSTRACT : Even with the establishment of Eurocode 7 in the European Union and its accompanying national application documents and standards for geotechnical investigation and testing there are still outstanding problems with the various national practices of pile design. This paper presents a revised interpretation of the prediction exercise for pile bearing capacity and settlement originally presented at ISP5 in August 2005. The field tests with the Ménard pressuremeter, the static cone penetrometer, and the SPT performed at the site, the answers to the exercise and a detailed investigation of the practice are discussed. The results are compared to the predictions of other national codes and of customarily used design methods and to the results from an instrumented CFA drilled shaft tested on the site. INTRODUCTION Determining the working load of a pile so as to be close to its actual bearing capacity is still very difficult. In most countries, development of piling techniques were made simultaneously with the establishment of specific design methods. At the same time, significant efforts have been made to improve the soil investigation on which these methods were based. France has a large variety of extremely complex soils for which intact sampling is not possible. Engineers have therefore favoured insitu testing. Their choice lay between the static cone penetration test (CPT) and Ménard pressuremeter tests (PMT). For a pile the static capacity Qu is computed from: Qu = Qpu + Qsu (1) where: Qpu is the ultimate pile toe (base) capacity and Qsu the ultimate skin (shaft) resistance capacity. This separation of the pile capacity into two terms is a common feature of all the design methods used in practice: analytical methods based on friction (proportional to I’-c') and empirical methods based on in situ tests (CPT, SPT, PMT). The tip capacity is related to a mean value of the shear strength deduced from laboratory or in situ tests multiplied by a factor related to the failure mechanism and adjusted for the soil type and for the remoulding effect of the installation technique. The shaft term accounts for the change of soil properties in the vicinity of the pile after it has been installed, for the soil variability and for the (complex) pile-soil interaction. For laboratory as for in situ tests, due to technical features such as penetration speed for CPT, length of the sampler for SPT or of the probe for PMT, the segmentation of the soil log varies. Hence, for every segment of the shaft, shaft resistance has to be computed from the shear strength times a factor depending of the three influences above.
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When PMT results are used, in the general case of a layered ground for which the distribution of pressuremeter limit pressures pLM with depth are known, each of these terms will be calculated from the following equations: Qpu = [qo+k.(pLMe-po)]SB²/4 (2) (3) Qsu =6in qsi.SB.li where: qo is the total vertical pressure, k is the bearing factor, po is the horizontal total pressure, B is the diameter of the pile, qs is the limiting unit shaft friction of the ith layer, l the thickness of the ith layer. pLMe the equivalent limit pressure defined as the geometric mean of the pLM values obtained near the tip of the pile. While it is unnecessary to comment on most of the parameters in these equations, it is important to appreciate how to determine the bearing capacity factor k and the unit shaft friction qs. Altogether they characterize the Ménard direct design method. The tip bearing factor k is found from charts based on a range of soil classes and depends on the nature of soil, its density (known by pLM) and the installation process. The limit unit shaft friction qs is given by empirical charts of limit pressure pLM plotted against the nature and density (pLM) of the soil, the method of installation of the pile and the nature of the pile shaft (Bustamante et al. 2009). First proposed in the 1960s, the charts were obtained by analysis of a limited number of pile load tests. Twenty years later and after about 186 load tests on 88 instrumented piles, the database was used to readjust the method (Bustamante and Gianeselli 1981; Combarieu 1990). A new Design Code for bridge foundations was set up as Fascicule 62-V (MELT 1993). The method is proposed in an Appendix to Eurocode 7-2 (CEN 2006) and is still under improvement by a steering group of the French standardisation committee (Combarieu and Canépa 2007). Before analysing the results of the ISP5 exercise, we briefly outline the data given to the participants. ISP5 BENCHMARK EXERCISE Outline At the ISP5 symposium commemorating the 50th anniversary of the first Ménard pressuremeter patent, a benchmarking exercise was organised (Reiffsteck 2005). The purpose was to calculate the bearing capacity under axial load of a test pile drilled using a continuous flight auger. The pile, 0.5 m in diameter and 12 m deep (Fig. 1a) was installed on an experimental site located in Merville in Northern France on a former WWII airfield. The soil is silt overlying highly overconsolidated and fissured Flanders clay (similar to London clay). Raw data from three pressuremeter soundings, each of them involving 14 PMT, together with the pressure loss correction curves, were given to the participants who were asked to make their own interpretation of the tests and to quote the method
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used. Additional geotechnical data were also provided such as CPT and SPT profiles (Figs. 1b and 1c). The participants were asked to calculate (i) the bearing capacity of the pile, (ii) the settlement under a load equal to one third of the limit load and (iii) the settlement under a load of 500 kN. Methods based on the Ménard pressuremeter test results were to be favoured, but alternative approaches were to be accepted. 0 1 2 3 4 5 6 7 8 910
0
1.8m
2.4m
N SPT
qc (MPa)
silt
10 20 30 40 50
0
0 2
2
SPT 1 SPT 2
4
CPT1
4
CPT2 CPT3
0.5 m
8
6 depth (m)
clay
depth (m)
12 m
6
8
10
10
12
12
14
14
a) 16 b) 16 c) FIG. 1 (a) Pile and soil sketch (b) CPT sounding (c) SPT sounding Analysis of Ménard pressuremeter tests The readings of the 42 tests were provided to the participants but it was decided for clarity reasons not to give the table of the pressure and volume readings at 15 s and 30 s at each pressure hold. This point has caused comment relating to the potential scattering of the data due to the analysis method (Long 2008). Because from a graphical analysis of the diagram (p, 'V60/30) it was not possible to estimate the creep pressure pf, participants had to explore the area between the group of readings in the pseudo-elastic phase of the pressuremeter curve and the group of readings at large strains, since the Ménard modulus EM is obtained from the first group of readings (AFNOR 1991). Figures 2a and 3a show the mean curves of the three logs proposed by nine of the participants. To a first approximation, the curves seem to be parallel, i.e. the trend for each participant is systematically to underestimate either the modulus or the creep pressure (participant No.6) or in opposition to do the inverse (participants Nos.5 and 8). For the limit pressure, the mean value of the 9 participants answers is very close to the one observed by the LCPC team (Fig. 2b) which is also in the range defined by the standard deviation of the 9 answers.
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Ménard limit pressure pLM (MPa)
Ménard limit pressure p LM (MPa) 0
0,5
1
1,5
2
0
2,5
1
1,5
2
2,5
6 8 10
4 6 depth (m)
4
mean mean modified LCPC
2
1 2 3 4 5 6 7 8 9
2
depth (m)
0,5
0
0
8
10
12
12
14
14
a) 16 b) 16 FIG. 2 Ménard limit pressure pLM (a) participants answers (b) mean values For the Ménard pressuremeter modulus EM, two answers diverge markedly. These answers came from participants Nos. 4 and 7 and may be attributable to local practices. If we omit these two results, the modified mean then shows a tendency towards underestimation on the part of the participants when compared with the LCPC analysis (Fig. 3b). 0
Ménard Pressuremeter Modulus EM (MPa) 10 20 30 40 50 60
0
0
4
depth (m)
6 8
1 2 3 4 5 6 7 8 9
mean 2 4
LCPC
8
10
10
12
12
14
14
16
mean modified
6 depth (m)
2
Ménard Pressuremeter Modulus EM (MPa) 10 20 30 40 50 60
0
a) 16 FIG. 3 Ménard pressuremeter modulus EM (a) participants answers (b) mean values
b)
A very reassuring fact is that the Ménard limit pressure pLM which is the principal parameter used in the design method for computing ultimate shaft and toe capacity, see equations (2) and (3), is determined with good repeatability in this exercise. The mean error is under 24 % for the analysis made by the 9 participants and less than 20 % between curves of the 3 soundings originally analysed at LCPC. It can be
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compared with the 33 % relative error of qc observed on 7 soundings of 10 cm² cone penetration tests at the Bothkennar site and 23 % of the acceptable error for Class 1 application of the CPT European Standard on the qc profile presented by Long (2008). PILE LOAD TEST The pile was loaded axially and instrumented by LPC removable extensometers over its entire length (Bustamante and Jézéquel 1989). The limit load Qu is conventionally defined as the settlement at pile head equal to the higher of the two values: either 20 mm or B/10, which is here equal to 50 mm (Fig. 4a) (MELT 1993). Defining the total limit load Qu and using the LPC removable extensometers to find the limit point resistance Qpu permits the determination of the limit shaft friction Qsu (Fig. 4b). It gives Qu = Qpu + Qsu = 373 + 939 = 1312 kN for 50 mm of displacement and a creep load of Qc = 1000 kN. The settlement yo at the top for a 500 kN load is 1 mm and at one third of the total limit load it is 0.75 mm. 0
200
400
vertical load (kN) 600 800 1000
Qu = 1312 kN 1200
1400
0
0
500
load (kN) 1000
1500
0
H
5
2
15
G
Qc = 1000 kN
20
4 depth z (m)
settlement yo at 60' (mm)
10
25 30
F
6
E 8
D C
35
10
B A
40
12 45
Qu = 1312 kN
50
Qpu = 373 kN 14 a) FIG. 4 Test results
b)
PILE CAPACITY AND SETTLEMENT ESTIMATE Prediction of bearing capacity The bearing capacity values computed by the participants using PMT direct design method lie spread between 0.6 and 1.4 times that measured. On average, the participants have underestimated the measured value. The mean value of the normalised answers is equal to 0.89 with a standard deviation of V = 0.23 (Fig. 5). The values of the bearing capacity predicted by the I’-c’ method based on laboratory tests are slightly overestimated, but in a range similar to that obtained by the pressuremeter method. For the CPT method used by three of the participants,
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calculation of point resistance was made using Begemann or Meyerhof methods and shaft resistance using different proposals for the skin friction factor (D and O methods among others). The bearing capacities deduced from CPT methods are all conservative. The discrepancy in the values for the bearing capacity prediction seems to be at least partially due to the design method used: I’-c’, CPT SPT or PMT.
Qu predicted/Qu observed
1,6
fI, c PMT CPT NUM SPT
1,4 1,2 1 0,8 0,6 0,4 0,2 0 1
2
3
4
5
6
7
8
9
10
11
participants
FIG. 5 Prediction of bearing capacity of ISP5 benchmark Table 1 shows a comparison of several design methods either proposed in Eurocode 7-2, or by national rules or, still, customarily used. The pressuremeter design method using the original LCPC log for pLM values slightly underpredicts the experimental bearing capacity of the CFA pile. In the calculation of bearing capacity using the Dutch CPT design method, where the three CPT soundings of fig. 2b are considered, overestimation is observed (NEN 1991). For the French national rule (Fascicule 62-V), the CPT based method, a large underestimation is seen. The D, E and O methods all underestimate the bearing capacity of the pile (Bowles 1997). Table No. 1 Comparisons Between Some Design Methods Used to Derive the Bearing Capacity in MN F62-V F62-V NEN 6743 B&2G D O E EC7-2 E EC7-2 D PMT CPT CPT PMT I’-c’ I’-c’ I’-c’ Qpu 0.373 0.346 0.227 0.343 0.254 0.168 0.168 0.168 Qsu 0.939 0.901 0.754 1.117 0.848 0.906 0.968 0.604 Qu 1.312 1.248 0.982 1.461 1.103 1.074 1.136 0.772 F62-V: Fascicule 62-V; EC7-2D: EN 1997-2 Annex D; EC7-2E: EN 1997-2 Annex E, B&2G: Bustamante et al., 2009. Note that 1MN =100 metric tonnes Q Load (MN) test
Prediction of settlement At present, no settlement prediction method is suggested for piles in Eurocode 7 part 2 (CEN 2008). In Fascicule 62-V (MELT 1993), two methods are proposed. In the first the settlement is arbitrarily defined as a percentage of the pile diameter
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(marked “A” in Fig. 6a). The second method is a determination of load transfer q-z curves as a function of Ménard pressuremeter modulus (marked “t-z” on Fig. 6a) as proposed by Frank and Zhao in 1982 (Gambin and Frank 2009). When input parameters come from the CPT, the method proposed by NEN 6743 (NEN 1991) relies on the same principle but q-z curves are smoothed curves. In addition to the elastic or semi-empirical methods, numerical methods have been used by three participants (“NUM” in Fig. 6a). Several methods have been applied in order to identify the constitutive parameters of the soil model adopted: laboratory test results given in the paper (participant No.1), parameters imposed in the benchmark data (participant No.1), correlations based on the Ménard pressuremeter modulus (participants Nos.8 and 9), inverse analysis of pressuremeter curves (participants Nos.6 and 10). Figure 6b shows the proposed load-settlement curves as calculated according to Fascicule 62-V and to NEN 6743 and as submitted by five participants in the benchmarking compared with the actual pile load test curve. Among conventional methods, it seems that the q-z or t-z curve method used by four of the participants gives results closest to reality (participants Nos.8 and 9 on fig. 6b). 8
0
A
vertical load (kN) 600 800 1000 1200 1400
t-z
10
NUM
5 NUM
NUM
3
A
t-z
A
2 t-z
1
t-z
settlement (mm)
yo predicted/yo observed
400
5
6
4
200
0
7
15 20 25
1
2
3
4
5
6
7
8
9 10 11 participants
1 6 8
30 35 40
0
Load test
9 10 NEN6743
45
a) 50 FIG. 6 Settlement prediction
b)
Table No. 2 Settlement yo prediction at Q=500 kN
Parameter origin Parameter values yo (mm)
Load Fasc. 62-V NEN Poulos &Davis (1974) test Frank & Zhao 6743 PMT CPT Triaxial Particip. No.9 table 1 Es=50 MPa 1 1 3.9 1.2
Bowles (1997) SPT N=25 1.6
Under a 500 kN load, the settlements estimated by PMT and CPT methods and more classical methods are very similar (Fig. 6a & table 2). It is difficult to judge the accuracy of these computations as the measured settlement is very small. During the benchmarking, the participants have overestimated the settlement by an average ratio of 3.42.
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CONCLUSIONS In the ISP5 pile prediction exercise, whichever the method used, the bearing capacity and settlements estimated were conservative compared to the actual pile load test results. Predicted bearing capacities have proven relatively close to observation while some of the settlements were seriously overestimated. This finding may be due to the nature of soil: a very overconsolidated and fissured clay leads to a difficult assessment of the mechanical characteristics. However the “t-z curve” method seems to give the best results. REFERENCES AFNOR (1991). Essai pressiométrique Ménard French Standard NF P 94-110, La Plaine Saint-Denis, 43 p. Bowles J.E. (1997) Foundation analysis and design, 5th edition, Mc Graw-Hill Int. Eds., 1175 p. Bustamante M., Gianeselli L. (1981). Prévision de la capacité portante des pieux isolés sous charge verticale, Règles pressiométriques et pénétrométriques, Bull. des Laboratoires des Ponts et Chaussées, 113: 83-108. Bustamante M., Jézéquel J.-F. (1989). Essai statique de pieu isolé sous charge axiale, Méthode d’essai LPC n°31, LCPC Paris, 12 p. Bustamante M., Gambin M., Gianeselli L. (2009). Pile Design at Failure Using the Ménard Pressuremeter : an Up-Date, Proc. IFCEE ’09, ASCE. Combarieu O. (1990). Comparaison des règles pressiométriques 1972 et 1985 de calcul de la capacité portante des pieux, Bull. des Laboratoires des Ponts et Chaussées, 170: 101-111. Combarieu O., Canépa Y.(2007). Méthode de calcul de la capacité portante des pieux, Working document– CNOG P94-262, 4 p. CEN (2006). Eurocode 7: Geotechnical design — Part 2: Ground investigation and testing, European Standard EN 1997-2, 222 p. Gambin M., Frank R. (2009). Direct Design Rules for Piles Using Ménard Pressuremeter, Proc. IFCEE ’09, ASCE. Long M. (2008). Design parameters from in situ tests in soft ground – recent developments, ISC’3 Taiwan, Geotechnical and Geophysical Site Characterization – Huang & Mayne (eds) Taylor & Francis Group, London, pp.89-116 MELT (1993). Règles techniques de calcul et de conception des fondations des ouvrages de génie civil, CCTG Fascicule 62 Titre V, Ministère de l’Équipement, du Logement et des Transports, Paris, Texte officiel N° 93-3, 182 p. NEN (1991). Calculation method for bearing capacity of pile foundation, compression pile, Dutch Standard NEN 6743, 31 p. Poulos H. G., Davis E. H. (1974). Elastic solutions for soil and rock mechanics, John Wyley & Sons, 411 p. Reiffsteck P. (2005). Portance et tassements des fondations profondes : présentation des résultats du concours de prévision, Symp. Int. ISP5-PRESSIO 2005, 50 ans de pressiomètres, Gambin et al. (eds.), Presses de l'ENPC/LCPC, 2 :521-536.
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Rades Bridge Drilled Shafts Designed and Tested Using Menard Pressuremeter François Schlosser1, Alain Guilloux2, Kamel Zaghouani3, Patrick Berthelot4. 1
Professor Em., Ecole Nat. des Ponts et Chaussées. Marne-la-Vallée, France ;
[email protected] General Manager, Terrasol, 72 av. Pasteur, 93108 Montreuil Cedex, France ;
[email protected] 3 Manager, Terrasol Tunisie, 2 rue M. Abdessalem, El Menzeh V, Tunis ;
[email protected] 4 Geotechnical Director, Veritas, 92400 Courbevoie, France;
[email protected] 2
ABSTRACT: The founding soil of the Rades-La Goulette cable-stayed bridge in the Tunis Lake is made up of compressible clays for depths exceeding 120 m. The two main towers were founded on groups of 2m OD and 80 m deep drilled shafts. Two soil surveys were carried out with the Menard pressuremeter down to 105 m depth and were used to predict the bearing capacity of the piles and to estimate their settlement. To check the predictions loading tests were performed, in particular a Class A one on a 1 meter OD and 80 meter deep test drilled shaft. Comparing the observed settlement versus load curve with the predicted one showed a fairly good agreement. INTRODUCTION The Rades-La Goulette cable-stayed bridge, constructed in the lake of Tunis area provides a north to south connection which avoids the Downtown area by spanning the Straight Canal linking the old Tunis harbor to the sea. The bridge is connected to the Tunis-La Goulette expressway by an interchange which has replaced the road along the old canal. Land reclaimed from the lake was required to construct the interchange for which embankments and access bridges were erected. The main part of the bridge consists of three cable-stayed spans, respectively 70 m, 120 m and 70 m long so providing a 20 m high and 70 m wide passage for ships. The two towers reach 40 m above sea level. Figure 1 shows a longitudinal section of the cable stayed bridge with the soil profile also shown. One of the main features of the project was the founding soil of compressible clays exceeding 120 meters deep. For this reason the foundations of the two main piers P12 and P13 was originally designed with a square group of 8 piles 2 m OD and 100 m deep. Shaft drilling started in 2005 with bentonite slurry using conventional auger boring down to 60 m depth and then reverse circulation drilling (RCD) below that. RCD was required to correct departures from verticality. The construction of the first drilled shaft for pier P12 faced difficulties with RCD due to a sticky clay from 60-70 m depth and it refused at 80 m depth. A new foundation was therefore designed for the two piers P12 and P13, consisting in a square group of 9 shorter piles at 75 m deep. Furthermore a loading test on a 1meter OD pile was commissioned in addition to an Osterberg pile load test performed in 2005. However the two adjacent piers P11 and P14 at the ends of the bridge remain founded on a square
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group of 8 piles, 1.5 m OD and 60 m deep.
FIG. 1. Longitudinal section of the cable stayed bridge and foundation. GEOTECHNICAL SURVEYS Two geotechnical surveys, one for the preliminary design and one for the construction design were required for the foundations of the main cable-stayed span, the access bridges and the interchange. As the Ménard pressuremeter (PMT) was considered to be a very useful tool for the design of deep foundations in compressible clays the first survey for the main span involved two pressuremeter boreholes down to 105 m depth and 2 boreholes to the same depth for soil sampling. The second survey involved 4 PMT soundings down to 115 m depth, 4 boreholes to the same depth for soil sampling and 2 CPTu (piezocone) soundings though these last refused at 25 m depth. Disturbed and undisturbed samples were used for determining index properties and for laboratory direct shear, triaxial and consolidation tests. The soil stratification in the cable-stayed bridge area was as follow: x mud and very soft clay (0.5 to 8.5 m depth) : layer I x sand and clayey sand (8.5 to 17.5 m depth) : layer II x plastic clay (17.5 to 24 m depth) : layer III x fine sand and fine silty sand (24 to 35.5m): layer IV x plastic medium stiff and stiff clay (35.5 m to 115 m) : layer V These five layers were divided into sub-layers as shown in figure 2 once the Ménard parameters were known. Fig. 2 also shows the pLM*(= pLM – p0) graphs for the six PMT soundings. It is interesting to note that the mean pLM* values in the clay layer V changes fairly linearly from 1 MPa at 35.5 m depth to 2.3 MPa at 90 m depth.
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FIG. 2. Soil parameter values from Ménard pressuremeter tests (PMT). DESIGN The PMT data were used to design the deep foundations of the cable-stayed and access spans both in the preliminary design and in the construction design. The bearing capacity and the settlement of the pile groups were estimated. However, only the construction design of the deep foundations for main piers P12 and P13 and for the adjacent piers P11 and P14, are presented here. Bearing capacity of the isolated shaft and of the shaft group The conventional Menard formula : ql – qo = kp.pLM* was used for estimating the limit load Qp at the isolated shaft tip or at a shaft group base, where the bearing factor kp as a function of soil type and pile type is obtained from the French Code of Practice (M.E.L.T. 1991; Frank 1994). The qs(z) values for estimating the total limiting friction resistance Qs along the shaft were also determined using this Code of Practice from qs, and pLM* curves plotted as a function of soil type and pile type. This method is based on the analysis of a large number of pile load tests (Bustamante et al. 2009). For an isolated drilled shaft 75 m deep and 2 m OD at piers P12 or P13 and a similar one 60 m deep and 1 m OD at piers P11 or P14, the limit bearing capacity at the tip Qp and in friction along the shaft Qs are given by equations (1) where kp = 1.2 : Qp = (kp.pLM* + q0).S
Qs ʌ'³0L qs(z).dz
(1)
From the French Code of Practice, the creep load Qc and the maximum serviceable load Qa of the shaft are given by equations (2) : Qc = 0.5 Qp + 0.7 Qs
Qa = Qc/1.4
(2)
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Numerical results are presented in table 1. The maximum serviceable load on the shaft must be checked against the (Q + Wp) load where Q is the maximum quasi permanent in-service load on the isolated shaft and Wp is the self-weight of the pile. The ratios for piers P11, P12, P13 and P14 are respectively 1.16, 1.05, 1.05 and 1.39, i.e. all greater than 1.0 as required. Table 1. Pile Tip Resistance Qp, Shaft resistance Qs, Creep Load Qc & Service Load Qa Shaft
Qp (MN)
Qs(MN)
Qc (MN)
3.92 7.99
12.25 20.36
10.54 18.25
P11 or P14 P12 or P13
Qa=Qc/1.4 (MN) 7.53 13.03
For the drilled pile group at piers P11, P12, P13 and P14, two methods were used for estimating the admissible load: (a) using the group effect efficiency coefficient Ce and (b) using the equivalent pile concept of a vertical cylinder enveloping the shafts. Both give acceptable values of the ratio R of the maximum serviceable load Qga to the actual service load (Q + W). R must, obviously, be greater than 1.0. However, method (a) gives smaller values of R than the equivalent pile method (b). Settlement prediction The pressuremeter modulus EM gave valuable settlement information at the deep foundations of all three bridges. The pile group at pier P12 is now presented. A preliminary calculation was performed for an isolated drilled shaft, 75 m deep and 2 m OD, subject to the maximum service load in one of the 9 pile group. The method of Frank and Zhao (1982) was used to find the mobilized shDIWIULFWLRQVWUHVVIJYHUVXVWKH local shaft displacement s, and the mobilized vertical stress q versus the tip displacement sp $V LQGLFDWHG LQ ILJXUH WKH FRUUHVSRQGLQJ FXUYHV IJ V DQG T Vp) have a non linear shape and involve 4 parameters : the maximum values qs, qp and two gradients kIJ and kq.
),*)UDQNDQG=KDR VKDIWIULFWLRQIJDQGWLSUHVLVWDQFHTLQSLOHV
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For soil and reinforced concrete, a short term and a long term E-modulus were considered. A pressuremeter modulus at long term was set at 1.5 times the short term value. The reinforced concrete modulus values calculated are respectively 40 600 MPa in the short term and 22 000 MPa in the long term, taking into account the steel reinforcement effect. The distribution of shear stress IJ ] DQGRIORDG4 ] DUHGHWHUPLQHGIURPGLIIHUHQWLDO equations resulting from considerations of a) local equilibrium of the shaft; b) linear elasticity of the shaft; c) Frank and Zhao mobilization of qs and qu stresses. A settlement calculation of the isolated pile of pier P12 was then performed using two methods: (a) Frank and Zhao method using the FOXTA computer program of Simon et al., (2002) and (b) the Finite Element Method (FEM). using PLAXIS and an elastoplastic behavior of the soil in which the elastic modulus E is taken as EM Į 7KH settlement values for the maximum in-service load are as follows : where sst : short term settlement FOXTA : sst = 5.3 mm; slt = 8.9 mm slt : long term settlement PLAXIS : sst = 14 mm; slt = 26 mm Settlements calculated by PLAXIS are around 3 times those calculated by FOXTA. However, the Frank and Zhao method gives a settlement fairly close to the actual value as will be seen later. In order to get similar results with PLAXIS, the elastic modulus of the soil would have to be taken as equal to 3 EM /Į)LJXUHVKRZVWKHORDG-settlement curve calculated by FOXTA for the isolated pile above. There is a sudden change in the slope at 19 MN pile head load. This corresponds fairly well to the creep load Qc calculated as 18.25 MN.
FIG. 4 Calculated load-settlement curve for the 75 m long, 2 m OD pile. The settlement of the group of 9 piles under pier P12 was calculated by the same two methods as for the isolated pile. The group was also modeled by an equivalent pile enveloping the drilled shafts and the soil between them, i.e. a pile section close to a square assumed to behave as a solid body. The equivalent elastic modulus of this pile was calculated from the elastic modulus of the drilled shaft and of the soil with E = EM Į/LPLWLQJVKDIWIULFWLRQTs of the soil elements along the equivalent pile were taken as equal to the soil shear strength and the same value of the bearing factor at the base kp = 1.2 was taken as for isolated pile. The FEM calculation was in two dimensions assuming axisymmetry. Settlement values are as follows:
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FOXTA : sst = 13.6 mm; slt = 21.3 mm where sst : short term settlement slt = 52 mm slt : long term settlement PLAXIS : sst = 37 mm; As for the isolated drilled shaft, the settlements calculated by FEM are roughly 3 times those calculated by Frank and Zhao, which are in the range of the observed E/EM values (Baguelin et al. 1986). This leads to a stiffness of the drilled shaft group under the service load about 3 times smaller than that of one of the 9 isolated piles. It results from the fact that shaft friction is proportionally smaller in the group than in the isolated pile. Calculations from the FOXTA program using Frank and Zhao were considered to be the most representative ones. PILE LOADING TESTS Österberg pile load test The first pile loading test using the Osterberg O-cell procedure was performed in 2005. It consists of loading a pile constructed in two parts approximately 2/3 and 1/3 of the total length and with a horizontal hydraulic jack located at the junction. Strain gauges are placed at the tip for measuring the mobilization of the tip resistance .This particular pile was 1.5 m OD and 63 m long. Figure 5 shows a load–settlement curve obtained from the test data of a conventional load test. Also shown is the curve calculated by the Frank and Zhao method using the FOXTA program. The two curves coincide until the load attains the creep load Qc of 10 MN for the two curves. After creep load, failure is much more rapidly reached in the experimental curve. This difference was identified as a defect at the drilled shaft tip and, it was decided as a result to grout all drilled shaft tips. Instrumented pile test with loading at the pile head To show that a reduction of the drilled shaft length to 75 m for the foundations of piers P12 and P13 was possible a conventional Class A pile loading test was performed according to the French Standard. The pile was equipped with vibrating wire strain gauges to measure the distribution of the load along the pile. The pile tested was 80 m deep and 1.5 m in diameter. Figure 6 shows the experimental load-settlement (Q, s) curve and the one predicted by FOXTA using the Frank-Zhao method. Contrary to the Osterberg pile load test results, the comparison between experimental and predicted curves shows a good agreement for the limit load. QLE= 12.5 MN for a B/10 settlement and the predicted QL = 12.1 MN. However, some differences in the creep load, experimental Qc = 10.2 MN and predicted Qc = 8.1 MN, as well as in the pile stiffness which was observed to be twice the predicted value. This observation increased the safety of the bridge foundations and the results of this second pile load test justified the use of reduced length piles in the final construction.
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FIG. 5. Conventional load-settlement curves from results of the Österberg pile load test with the predicted curve for a pile 60 m deep and 1.5 m OD. FIG. 6. Measured and calculated loadsettlement curves for the class A drilled shaft loading test (pile: 80 m length, 1 m in diameter) It is interesting to look at graphs of the load distribution in the pile as presented in figure 7.
FIG. 7. (Q,z) distribution during loading of the 1m OD drilled shaft. In each Q-z curve, the slope is proportional to the value of the mobilized shaft friction IJDWWKLVOHYHO ǻ4ǻ] 6IJZKHUH6LVWKHSLOHFURVVVHFWLRQDODUHD)RUWKHORDGDWWKH pile head Q0 = 1.5 MN the shaft friction is only slightly mobilized and there is practically no load in the shaft below 55 m depth. Under increasing load, the whole pile is progressively stressed until the whole shaft friction is mobilized. This corresponds to Q0 = 9 MN where the tip resistance is not yet mobilized and the shaft behaves as a
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friction pile. As Q0 increases the tip resistance is progressively mobilized and the load distribution graph translates upwards. Since the creep load Qc is 10.2 MN, the pile behaves as a friction pile for all Q0 values in the range (0, Qa) of the shaft service loads, where Qa = Qc/1.4 = 7.3 MN. CONCLUSION The Menard pressuremeter was a most valuable tool for the design of very deep foundations in compressible clays at the Rades bridge. It easily characterized the soil down to 105 meters deep and reliably measured important soil properties. It is shown here to be very well adapted to the direct design of deep foundations both for bearing capacity and for settlement. This finding confirms the results of large numbers of instrumented pile load tests. The fairly good agreement between estimated and measured settlements in the instrumented pile loading tests performed for the Rades bridge is an additional proof of its capability. ACKNOWLEDGMENTS The authors are grateful to the Ministry of Public Works of Tunisia and to Taïsei Corporation for permitting to use the data of the investigations and the pile loading tests. REFERENCES Baguelin, F., Bustamante, M., Frank, R. (1986). “The pressuremeter for foundations: French experience”. Proc. Conference on the Use of In-Situ Tests in Geotechnical Engineering, Blacksburg, VA, ASCE, Geot. Special Pub., No.6, pp.31-46. Bustamante, M., Gambin, M., Gianeselli, L. (2009) Pile Design at Failure Using the Ménard Pressuremeter: an Up-Date, Proc. IFCEE ’09, ASCE. Frank, R., Zhao S.R. (1982). Estimation par les paramètres pressiométriques de l’enfoncement sous charge axiale des pieux forés dans les sols fins. Bull. Liaison Labo P. et Ch., n° 119 : 17-24. Frank, R. (1994). The new Eurocode and the new French code for the design of the deep foundations. Proc. Int. Conf. Design and Construction of Deep Foundations. Orlando. Florida. FHWA Vol.1: 279-304. M.E.L.T. (1991). Design Rules for Foundations, Tender Documents for Public Works. CCTG, Fasc. n° 62, TitreV (in French). Imprimerie Nationale Paris. Simon, B., Kazmierczak, J.B., Bernhardt, V. (2002). Benefits from a modular foundation design software. Proc. 5th Eur. Conf. Numerical Methods in Geot. Eng. NUMGE, Mestat (ed.) 2002, Presses de l'ENPC/LCPC, Paris; 357 – 362.
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LNG Tanks at Damietta on Drilled Shafts Designed and Tested Using Ménard PMT Jean-Yves Boumedi1, Jean-Pierre Baud2 & Bruno Radiguet3 1
Department Head, Bouygues-Travaux Publics, Le Challenger, St. Quentin en Yvelines, France,
[email protected] CEO, Eurogeo, Avrainville, France,
[email protected] 3 Head, Design Office, Bouygues-Travaux Publics, Le Challenger, St. Quentin en Yvelines, France,
[email protected] 2
ABSTRACT: The Damietta Liquefied Natural Gas (LNG) tank farm is situated in the lower Nile Valley, in an area prone to earthquakes. The design of the piles was based on a combination of various in situ surveys, using mainly Ménard PMT. The results of vertical and horizontal load tests showed good agreement with predictions based on Ménard PMT direct design methods. 1. INTRODUCTION The gas liquefaction plant of Damietta (Dumyāt) in Egypt is located along the eastern stream of the Nile delta. It will produce 5 million tonnes of LNG per year. It includes 4 tanks 40 m high, 80 m in diameter and on 155 m centers, each is designed to store 150 000 m3 of liquefied gas. The storage tank design integrates a pre-stressed concrete wall with an above-ground dual metal shell structure. The two Train # 1 reservoirs, built in 2003 (Union Fenosa 2006) are the subject of this paper. 2. GEOLOGY Damietta Harbour, located on the Mediterranean, is in a seismic area due to the colliding of the African and Eurasian tectonic plates. The zone is classified 3 on the Egyptian Seismic Scale. Tsunamis were recorded in the years 365, 1303 and 1908 (Riad and Yousef 1999). The Nile delta recent deposits include sub-horizontal layers of gravel, sand, silt and clay interbedded over depths exceeding 1000 m. These deposits date from the end of the Tertiary era to the present day. Several aquifers fed by the Nile river flow through this geological profile. Several site investigations were carried out at various stages of the project, including two boreholes down to 45 and 95 m depth with core-sampling and SPT’s, 15 CPTu’s
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down to 40 m around the proposed tanks and in their center, though many of these refused at 25 m, and three Ménard PMT (ASTM D 4719) soundings to 65 and 95 m depth at each tank location. The soil profile is as follows : - 0 – 12 m the upper sand layer (USL) - 12 – 28 m black soft clay (SoC) - 28 – 60 m yellowish fine to medium sand underlying an intermittent layer of stiff silty clay (DS & SC) - 60 – 95 m grey to black stiff clay with peat pockets (HC) - 95 – 110 m dense sands (DS)
FIG. 1. Ménard PMT data: (a) Pressuremeter Modulus EM, and (b) Limit Pressure pLM. Ménard PMT data logs are plotted together in Fig.1 and mean values for the 6 boreholes are listed below in Table 1 :
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Table 1. Mean Ménard PMT Data per Layer Layer Upper Sand Soft Clay Stiff Silty Clay Dense Sand Hard Clay
Upper boundary (m) +1.4 -12 -29 -36 -60
EM (MPa) 5.8 8.1 12.2 20.7 22.1
pLM (MPa) 0.69 0.57 1.55 2.98 2.59
3. TANK FOUNDATIONS Each tank applies a pressure of 0.264 MPa to the soil. The foundation is resting on 1.20 m (4 ft) O.D. drilled shafts, 316 in number and 47 m deep. Together they represent 15 300 m of large diameter drilling and 20 000 m3 of concrete.
FIG.2 Ménard limit pressure under the tanks. Each shaft, drilled using bentonite slurry, required 7.4 metric tonnes of reinforcement and 53 m3 of concrete. Sonic control pipes were installed in 48 of the piles. One vertical load test and one lateral load test were performed, each following ASTM D1143:1 and ASTM D3966:1. In addition, 29 piles were subjected to sonic testing, all 316 piles were tested by the impedance method and two were dynamically tested.
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4. PILE LOAD TESTS Vertical Load Testing The vertical test pile was located between the 2 tanks as in Fig. 2 and was drilled down to a depth of 47.80 m at a diameter of 1200 mm. It was loaded up to 1000 tons (Fig. 3), i.e. twice the service capacity of the foundation piles. TIME - LOAD DIAGRAM
TIME - SETTLEMENTS DIAGRAM
1200
Time (minutes) 0
800
Settlement (mm)
Load (tonnes)
1000
600
Pile diameter 120 cm Pile length : 47,80 m
400 200 0 0
500
1000
1500
2000
2500
3000
Time (minutes)
500
1000
1500
2000
2500
3000
0 1 2 3 4 5 6 7 8 9 10
FIG. 3 Vertical load testing data. Ultimate Bearing Capacity Various methods were used to assess the ultimate pile capacity, among them the Chin modified method (Chin 1978) in which the ratio of settlement divided by corresponding applied load is plotted against the settlement (Fig. 4). The slope of the best linear fit to the graph gives 1364 tonnes here corresponding to a pile stress q =11.8 MPa. This method has been recently tested on 50 instrumented piles in various soils (Borel et al. 2000). The authors there showed that if the test is not carried close enough to failure, the assumptions lead to an underestimate of pile ultimate capacity. CHIN'S MODIFIED METHOD (1978)
LOAD - SETTLEMENT DIAGRAM
200
400
600
800
1000
1200
tonnes)
Settlement (mm)
0 0 1 2 3 4 5 6 7 8 9 10
Settlements / Load (mm /
Load (tonnes)
0.01 0.009 0.008 0.007 0.006 0.005 0.004 0.003 0.002 0.001 0
y = 6.38E-04x + 3.83E-03
0
2
4
6
Settlement (mm)
FIG. 4 Vertical loading test curve and Chin method extrapolation.
8
10
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According to EUROCODE 7, part 2 (CEN 2006), the ultimate load Qult can be calculated from Ménard pressuremeter tests using the equation : Qult = Qpu + Qsu
(1)
(2) - the ultimate tip load being Qpu = A x kp x [pLM – p0] where A is the pile section, kp a bearing factor based on soil type and p0 is the horizontal earth pressure at rest, (Gambin & Frank 2009), - and the ultimate skin friction Qsu = P x 6 [qsi x zi] where P is the pile perimeter and qsi the unit shaft resistance at a depth zi.
(3)
By using these equations and the Ménard PMT design parameters, listed in Table 2, it is possible to predict Qult = 1554 tonnes (qult = 13.48 MPa) which is rather higher than Qult found by extrapolating the vertical pile loading by Chin's method. Table 2. Ménard Design Parameters by Layer Layer Upper Sand Soft Clay Stiff Silty Clay Dense Sand
pLM (MPa) 0.70 0.57 1.55 3.00
kp 1.2
qsi (MPa) 0.04 0.03 0.10 0.14
zi (m) 14.0 17.0 8.0 24.0
According to the French Design Code (M.E.L.T. 1993), the creep bearing capacity Qc corresponding to the end of the pseudo elastic resistance for the pile, can be obtained by Qc = 0.5 Qpu + 0.7 Qsu
(4)
Here, the Qc value was 985 tonnes (q = 8.5 MPa), which was conservative. Pile Settlement Estimate An iterative method to estimate pile settlement from Menard pressuremeter data was proposed in the early years of the development of this technique (Gambin 1963). In this method a small displacement w at the pile tip under a first load step is assumed and which results in a small tip pressure q , given by : w = (q / 2EM) . 0.3 (λB/0.6)
(5)
where O is a shape factor, The skin friction then mobilized in each soil layer for a displacement wi is calculated by :
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α
wi = Cj (qsi / EM) . 0.3 (R/0.6) (6) where Cj varies with the type of pile, qsi is the mobilized skin friction for the displacement wi limited by a value now given by EUROCODE 7 (CEN 2006), α is the rheological factor of the layer exhibiting a modulus EM. Good agreement with pile test data was obtained using a coefficient Cj = 1.1 (Fig. 5). 0
2
4
6
8
10
12
0
Stress q (MPa)
0.1 0.2 0.3 0.4
qc = 10.35 MPa
0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4 1.5
Loading test Settlement after Gambin (1963) Settlement after Frank & Zhao (1982) Settlement w (cm)
FIG. 5 Adjustment of calculation methods on loading test curve. Today, the French Design Code (MELT 1993) and EUROCODE 7 recommend the use of the more sophisticated method introduced for fine soils by Frank and Zhao (1982). The skin friction mobilisation equation becomes qsi = Kt . w up to qsi = qs /2 for that soil layer and then qsi = (Kt /5) w up to qs, (Gambin & Frank 2009), where the modulus of subgrade reaction Kt is given : - for fine soils : Kt = 2.0 EM/B - for granular soils : Kt = 0.8 EM/B This method was tested with the PIVER software of the French Government Laboratories for Bridges and Roads (LPC’s). The best fit was obtained for Kt = 1.3 EM/B in all fine sands and silt layers (Fig.5). For these two methods at this site, the authors think it necessary to use the pile parameters closer to “fine soils” or “clays” than to sands, due to the very fine grain size of sands and silts in all layers.
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Lateral Loading Test A lateral loading test between the same pile and an additional reaction pile was carried out according to ASTM D 3966.
Load (tonnes)
The applied lateral load was 100 tonnes, i.e. twice the design service load for the pile. The results (Fig.6) typically show a reaction with two linear segments corresponding to the two moduli of subgrade reaction, the second being half the first, as it appears on diagram in article 3.1 in Appendix C5 of the French Design Code for Public Works (MELT 1991), for short time action. 110 100 90 80 70 60 50 40 30 20 10 0
85 tonnes
K/2 50 tonnes
K 18,32 mm 5,01 mm
0
10
20
30
Displacement (mm)
FIG. 6 Results of the lateral load test. It was possible to calculate the value of this reaction from PMT data, using a "modulus" Kf derived from EM. Here Kf = 40 MPa. The maximum displacement G can then be estimated with pL = 0.7 MPa as measured in the top meter of the pile. The value obtained, G = 31 mm compares well with the observation of 28 mm measured at 100 tonnes loading. By following the original Ménard Direct Design Rules (Gambin 1979), the head displacement of a pile deeper than 3 times the elastic length L0 is given by : y0 = 2T0/L0.k.B
(7)
Here L0 = 5 m and the average pressuremeter modulus EM = 5 MPa. The estimated displacement under 85 tonnes is 16.4 mm, in good agreement with observation.
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5. CONCLUSIONS AND ACKNOWLEDGMENTS Union Fenosa LNG tanks at Damietta, founded on 316 piles, 47 m deep, were successfully designed using PMT data. The various loading tests carried out on a test pile proved the accuracy of the Menard Direct Design Rules. The influence of the soil grain size was pointed out. The authors appreciate the collaboration of Prof. R. Frank and J.-C. Dupla (Cermes, Marne-la-Vallée) for having let them use the PIVER software, and Mr. Clive Dalton (Cambridge Insitu Ltd, UK) for reviewing the final version of the text. 6. REFERENCES ASTM (2008). D 1143-1 Standard Test Method for Piles Under Axial Compressive Load, D 3966 Standard Test Method for Piles Under Lateral Load and D 4719 Standard Test Methods for Prebored Pressuremeter Testing in Soils, Annual Book of Standards. Borel, S., Bustamante, M., Gianeselli, L. (2004). “An appraisal of the Chin method based on 50 instrumented pile tests”, Ground Engineering, January, Vol.37, No.1, pp.22-26. Bustamante, M., Frank, R. (1999). “Current French Design Practice for Axially Loaded Piles” Ground Engineering, March, London, pp.38 – 44. Bustamante, M., Gambin, M. and Gianeselli, L. (2009). " Pile Design at Failure Using the Ménard Pressuremeter", Proc. IFCEE ’09, ASCE. CEN (2006). Eurocode 7 Geotechnical design – Part 2: Ground investigation and testing, pp. 123-124, Brussels. Chin F.K. (1978). Diagnosis of pile condition, Geotechnical Engineering, vol. 9, pp. 85104. Frank, R. and Zhao, S. R. (1982). "Estimation par les paramètres pressiométriques de l'enfoncement sous charge axiale des pieux forés dans les sols fins", Bull. liaison Labo P. et Ch., n° 119, mai-juin 1982, pp. 17-24. Gambin, M. (1963). "Settlement of a deep foundation in terms of PMT data" (in French), Sols-Soils, n°7, Paris, pp. 11-31 (large abstracts in English and in German). Gambin, M. (1979). Calculation of Foundations Subjected to Horizontal Forces Using Pressuremeter Data, Sols-Soils, No. 30-31, Paris, pp.17-59. Gambin, M. and Frank, R. (2009). "Direct Design Rules for Piles Using Ménard Pressuremeter, Proc. IFCEE ‘09, ASCE. M.E.L.T. (1993). Design Rules for Foundations, Tender Documents for Public Works. CCTG, Fascicule 62, Titre V (in French), Imprimerie Nationale, Paris Riad, S. and Yousef, M. (1999). Earth Hazards Assessment on the Southern Part of the Western Desert of Egypt, Assiut University Report, Egypt, January. Union Fenosa (2006). "Damietta LNG plant, first LNG plant in Egypt", GE Oil and Gas Annual Meeting, Florence, January 30-31, 2006 (available at www.unionfenosa.com).
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