EXPERIMENTAL AND NUMERICAL ASSESSMENT OF OSTERBERG LOAD TESTS ON LARGE BORED PILES IN SAND Roberto Nova, Politecnico di Milano, ITALY Bruno Becci, Ce.A.S. S.r.l., Milano, ITALY

The selection of the Osterberg Cell (O-Cell) technology as the standard testing method for deep foundations of the Railways Po Viaduct in Northern Italy offered a relevant number of field measures on large bored pile behaviour in alluvial soils. The assessment of pile response during such “non standard” testing procedure was performed through the comparison with standard top load tests on companion piles as well as by numerical models of the tests including a simplified pile-soil interaction scheme. As shown in this paper, the O-Cell technology allowed a careful assessment of the non-linear pile behaviour even at quite small loading levels as those required by a posteriori proof tests on production piles. Simple numerical models proved to be very effective in the simulation and the interpretation of pile behaviour during such unconventional testing procedure.

1. INTRODUCTION In the construction of the High Speed Railway Viaduct crossing the Po river near Piacenza (Italy), large diameter bored piles were routinely adopted as foundation system. High design loads and complex site conditions suggested the use of the Osterberg Cell (O-Cell) technology (Osterberg (1989)) as the first choice testing method for these piles. At onshore viaduct piers, usual kentledge load tests were also performed, thus permitting worthwhile comparison between different testing methods. All the performed tests were also reproduced through simple numerical models that offered a useful and thorough assessment of the measured pile behaviour during the O-Cell tests and a confirmation of the post-processing procedures employed in their interpretation. In the following, a review of the pile design criteria is reported; then a discussion of both design and proof load tests is included; numerical analyses of all the performed tests are presented and relevant parameters that best fit experimental observations are outlined. In the light of all these observations, some general remarks on the design of large diameter piles in sand are proposed.

2. SITE AND VIADUCT DESCRIPTION The Po Viaduct (Evangelista et al. (2003)) includes 23 bays, of which 20 approaching the river at both sides. A cable-stayed bridge, whose central bay is 192 m long, is placed at the permanent riverbed crossing (fig. 1).

Figure 1: the Po Viaduct near Piacenza (Italy)

Two meter diameter bored piles were used at the base of all the 24 piers, with pile lengths ranging from 40 to 62.5 m to ensure allowable loads between 10 to 18 MN. Bentonite slurry was employed in borings. Pile number per pier varies from 4 to 28.

Subsoil conditions at the site are represented by very deep submerged alluvial deposits: currently a shallow 15 to 20 m thick sand layer is overlying a 7 to 15 m thick silty clay stratum, which in turn is resting over a very deep silty sand deposit. Shallow and deep granular layers were mainly investigated by means of SPT and CPT tests that revealed almost normally consolidated sands; limited to offshore piers, very deep layers were sampled through a special CPT test, using self-penetrating equipment into advancing hole. In the granular deposits, SPT blow count N could be approximated by the following analytical expression: N=(0.78 to 1.0) · z (1) in which z is the depth in [m]; typical relative densities between 50% and 60% were estimated and an almost constant peak friction angle equal to 31° was evaluated, including dilatancy reduction due to high pressure, according to Bolton (1986). The intermediate, slightly overconsolidated, silty clay layer was analyzed through both in-situ CPT and lab tests, showing characteristic undrained shear strength values ranging from 50 to 100 kPa at depths between 15 and 30 m. This intermediate cohesive layer plays an important role in the hydraulic design of the offshore piers, as it limits the expected scour depth during extreme design flood conditions. 3. PILE DESIGN CRITERIA AND PRELIM-

INARY EXPERIMENTAL FINDINGS Preliminary design criteria are discussed in Becci et al. (2007). In Table 1 (left column) a summary of initial design assumptions is included.

Unit shaft resistance in sands was assumed linearly varying with depth through a constant coefficient K set to 0.6 and a friction parameter tan(δ): as for δ, the critical friction angle of soil, set to 30°, was selected. Such unit resistance was assumed to develop at a differential pile-soil local movement of 0.5% of pile diameter D. As for toe resistance, the design value in Table 1 was considered to develop corresponding with a toe displacement of 5% ·D. Before actual production pile construction stages, two design load tests were performed, according to the ASTM D1143 Quick Load Test Method, on sacrificial piles (50 and 55 m long), near actual Viaduct pier on the left hand riverside, i.e. at the north side of Po river. By means of a pair of Osterberg cells, installed in each shaft, as shown in fig. 2, ultimate loads could be almost reached.

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Figure 2: preliminary pile load test assembly

In both piles, the lower O-Cell was placed 2 m above the toe, thus allowing a close depiction of toe behaviour. Moreover, through the installation of ten strain gauge levels along the shaft, shear distribution at relevant test stages was obtained

Table 1

Preliminary Design assumptions Preliminary load tests results

Shaft resistance

Sands qs= K·tan(δ)· σ’v = 0.346·σ’v (2) Clay qs≅ 0 (3)

Sands qs= (0.62 to 0.85) ·σ’v (5) Clay qs≅ (0.20 to 0.25) ·σ’v (6)

Toe resistance Sands qb≅ N’q σ’v ≅10·σ’v (4) Sands qb≅ (7 to 9) ·σ’v (7)

σ’v =effective overburden stress = γ’·z γ’ = buoyancy unit weight= 10 kN/m3

by the measurement of the pile axial strain distribution. Based on these measures (Becci et al. (2007)), shaft and toe resistances listed in right column of Table 1 were estimated. It should be noted that a slightly lower toe resistance at 5%·D displacement was measured, whilst a higher shaft strength could be observed, the latter, however obtained through unusually large shaft displacements that could be imposed thanks to the particular testing method offered by the O-Cell technology. Increasing shaft resistance with depth was confirmed. The inspection of the equivalent top load curves obtained by an a posteriori processing of measures (solid gray curves in fig. 3) shows that preliminary design assumptions (dashed lines) had predicted a lower ultimate loading due to an underestimate of shaft resistance.

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Figure 3: Top Load vs Top Displacement curve for preliminary test piles

On the other hand, it was noted that the presumed pile stiffness at working load level had been well matched (see fig. 3, again); moreover, actual pile construction would have implied more difficult working conditions than those occurred during test pile installation; finally, long term extreme scour conditions for offshore piers would have represented an important issue that could have been hardly investigated by similar preliminary tests. For all these reasons, in spite of the apparently conservative preliminary design criteria, no pile

length was reduced in the final design. This choice demonstrated to be wise, in the light of actual construction observations. 4. NUMERICAL SIMULATION OF PREL-

IMINARY LOAD TESTS In the practice, the simulation of the soil-pile interaction by means of non-linear Winkler springs (O’Neill & Reese, (1999)) is still probably the most popular method to analyse single pile behaviour in inhomogeneous soil conditions. Adopting this approach within a finite element framework, various pile and loading conditions can be easily modelled, including the simulation of an Osterberg test as well. Using this method, the dark solid curves in fig. 3 have been computed, including strength and stiffness parameters summarized in fig. 4.

The toe reaction pattern has been included by assigning the behaviour as directly measured by O-Cells tests. At 5%·D toe displacement, the ratio N’q=qb/σ’v was found to be about 8.6 for Pile A and 6.7 for Pile B.

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Figure 4: numerical back-analysis on preliminary test piles

As for shaft reaction, trial shear strength profiles (solid lines in fig. 4) along with a threshold side displacement have been assigned. The latter is included via a scale factor η, which is actually one of the free parameters in this back-analysis process: note that, taking η=1, the normalized shaft displacement s/D as usually considered by O’Neill & Reese, (1999) is recovered. Final set of data, giving a good reproduction of target results, shows that η is currently higher than usual values; as for pile B, shaft reaction was found to be weaker yet stiffer than in pile A, in the deep sand layer. Improved numerical simulations can be obtained provided a slightly less smooth curve for shaft reaction in sand is adopted: this modification is not relevant for current discussion, however. 5. PROOF LOAD TESTS WITH OSTER-

BERG CELLS In preliminary load tests, due to high expected ultimate loadings, the O-Cell technology was deemed almost mandatory; however, proof load tests, up to 1.20 the maximum service pile load Nmax, could have been performed by means of traditional methods as well, unless quite complex conditions would have to be dealt with in offshore piers no. 7 and 8 (fig. 5).

Following some debate, three production piles in offshore piers and two onshore piles (fig. 6) were equipped with one O-Cell only, properly designed to impose a sufficient equivalent load for proof purpose. Near to the onshore O-Cell piles, three traditional top load tests were also requested.

Figure 5: site conditions for Piers 7 and 8

The O-cell was placed at about 20% of the pile length above the toe, as deep as possible to avoid significant lateral pile resistance reduction and, in the same time, to allow desired equivalent top load imposition without excessive cell loadings or toe movements.

Figure 6: piling layout and maximum load levels for proof load tests

Dealing with one O-Cell only offered the opportunity to perform a numerical analysis of the test itself, with a straightforward modification of the finite element model used to analyze the behaviour of the pile subjected to vertical top loads. To model the single O-Cell test, the finite element corresponding with cell position must be annealed and two equal loads must be applied in opposite directions to the nodes corresponding with annealed element. It should be noted that the simulation of a single O-Cell test requires no complex constitutive law for soil springs, since all such elements basically undergo monotonic loading path only. 6. PROOF LOAD TESTS DISCUSSION In the following, the results from each O-Cell test are discussed. Also these tests have been conducted according to the Quick Load Test Method. Limited to tests 1 to 3 (listed in fig. 6), remarkable nonlinear aspects have been highlighted during the tests; therefore some considerations about ultimate loading can be attempted as well. In onshore tests 4 and 5, nonlinear behaviour is less clear: on the other hand, a worthwhile comparison with traditional load tests is available, which can be considered a contribution to experimental assessment of O-Cell technology. OFFSHORE PIERS Fig. 7 outlines the results obtained for offshore piles in piers 7 and 8, by plotting the O-Cell plate absolute movements in upward (top plate) and downward (bottom plate) direction. In spite of applied (equivalent) loads slightly higher than maximum expected working loads, quite evident non-linear behaviour is obtained. Whereas such a behavior would, in general, be undesirable in a proof test, it is currently acceptable in such a procedure: in particular, a numerical simulation of each of these tests can sufficiently explain that the observed behaviour is physiological.

For back-analyses, the pile-soil interaction described in the previous section is used, with the normalized curves shown in fig. 8, same to those included in fig. 4, for shaft reaction.

In fig. 8 again, qs profile and relevant parameters obtained through the back-analyses are summarized; obtained responses are included in fig. 7, superimposed to experimental data. The back-calculated critical toe pressure qbLIM=4 MPa, corresponding with a toe settlement equal to 0.05 D is a good assumption for all the analyzed cases, and well agrees with eq. 7.

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Figure 7 – measured displacements and back analysis results for offshore tests

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In TEST 1 or 3, the shaft resistances qs and their mobilization levels quite well adhere to original design assumptions (as included in Table 1, left column), except that a non-zero shaft resistance is highlighted in the clay layer as well. As for TEST 2, the shaft behaviour above the O-Cell is almost the same as in the others; in the lower part, some reduced shaft resistances had to be included, to match the observed behaviour. At this stage, it is however hard to state whether this diminished response is due to locally weaker shaft resistance or to higher local compliance. To reliably depict the actual behaviour, the load would have had to be increased to a much higher level. The scaling parameter η decreases with qs: this numerical effect is necessary to reproduce similar side skin friction stiffness for all cases, independently from maximum strength value.

Anyhow, the back-figured parameters in fig. 8 currently fall within typical ranges. It should be noted that the shear strength increase with depth is confirmed: however, the mobilization levels tend to increase with depth as well, as also noted in the analysis of preliminary load tests discussed previously. ONSHORE PIERS In TEST 4 (fig. 9), both soil conditions and pile length are very similar to those considered for offshore piles. The measured behaviour was consistently very similar to TEST 1 or TEST3: in particular the TEST3 back-analysis can reasonably reproduce the behaviour of this pile too. Further modifications may be included, to better match higher stiffness of the upper part as well as a slightly lower stiffness of the lower segment: however such changes would not modify the overall description for this pile.

Finally, experimental results and numerical back-analysis of TEST 5 on the shortest pile in this campaign, is reported in fig. 10, whilst back-figured parameters are included in fig. 11. Back-figured shaft reaction displays a better response than in offshore piles: in sand layers, qs from eq. (1) can be increased by a factor 1.33, thus obtaining an average ratio β=qs/σ’v=0.46.

Figure 8: back analysis proof load test assumptions and results for offshore piles

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Figure 9 – measured displacements and back analysis results for onshore TEST 4

At adjacent pier no. 13 location (about 50 m far from pier 12, in S-E direction), corresponding with same piling layout and very similar soil properties, a traditional top load test was performed, using a steel kentledge. These results could be effectively compared with both equivalent top load curve obtained through the standard O-Cell post-processing procedure and numerical prediction based on back-figured parameters from TEST 5, in fig. 11. The comparison included in fig. 12 shows an excellent agreement among experimental data from traditional test (dots), equivalent O-Cell

curve (solid gray line) and numerical prediction (solid black line). The top load test could not investigate non-linear pile behaviour. Therefore, unless a top loading could be revised as a more natural way to test actual pile behaviour, the information offered by the traditional procedure is poor indeed as compared with an O-Cell test.

The remaining two top load tests were conducted at piers 11 and 18, on 46 m long piles. In both tests, an almost linear behaviour could be obtained, with top settlements of about 3 mm for both piles, at a proof load of 12 MN. 7. DISCUSSION AND CONCLUSIONS The selection of the Osterberg cell technology allowed the conduction of load tests to very high load levels that would have been hardly imposed due to complex environmental conditions. The O-Cell method in proof load tests could provide engineers with more useful and precise information than those currently available by traditional methods. In particular, some non-linear behaviour of part of the pile could be activated even at quite low loads. In addition to the routine post-processing of sampled data as a part of the standard O-Cell procedure, the authors performed simple numerical simulations of the tests using the Winkler method, and found that observed

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Figure 12 – comparison between traditional top load test, O-Cell test and numerical prediction for onshore pile

behaviour during these unconventional tests could be reasonably reproduced by means of interaction curves similar to those that would have been used in a traditional pile model. It is the authors’ opinion that these conclusions may contribute to increase the confidence by practicing engineers in the selection of modern testing techniques like the one discussed in this paper. It is important to realize that new testing methods currently allow the conduction of pile load tests, in almost all the real conditions. Complex site conditions and/or very high loads can, therefore, hardly be used as an excuse to limit or even omit load tests at all. On the other hand, a careful assessment of the obtained results is always recommended. In particular, heavy modifications to initial design, based on reasonable and well-established assumptions, should be considered with great care. As for the observed behaviour of these large diameter shafts, drilled under bentonite in sands, the authors found almost uniform toe behaviour, in good agreement with most frequently used correlations in the practice. As for shaft resistance, however, relevant discrepancies among design correlations, preliminary load tests and final proof load tests findings have been highlighted and discussed. These findings should, in general, suggest a particular care in the selection of shaft resistance parameters for bored piles in sand, all the more because similar observations have been reported by others, regarding bored shafts or barrettes in different soil conditions (e.g. Randolph (2003), Fellenius et al. (1999)). Such discrepancies are primarily related to construction issues, which can be hardly incorporated in preliminary design models (see Cernak (1976), Fleming & Sliwinski (1977), Ng & Lei (2003)). Even in the light of these limited observations, it can be argued that the suggested partial safety factors used in the growing LRFD method also in geotechnical engineering, may need some further discussion before being used in the practice.

In particular, reference is made to Eurocode 7 (CEN (2003)), that recommends a partial safety factor γB=1.60 for toe resistance, higher than the shaft resistance factor γS =1.30, for bored piles. The aforementioned values were likely tuned to implicitly account for different settlements necessary to activate each of the two contributions. However the authors argue whether such values may or may not conflict with some actual findings like those reported in this paper, as well as other frequent field observations in practical pile constructions. ACKNOWLEDGEMENT The technical advice of Grandi Lavori Fincosit staff, leaded by dr. Augusto Baù and dr. Raji Haykal, is particularly acknowledged. Piling Contractors TREVI S.p.A. and VIPP S.p.A., General Contractor SNAMPROGETTI and the Client Italferr S.p.A. are also acknowledged, as well as Loadtest Inc. engineers who provided and supported the Osterberg cell technology. REFERENCES BECCI, B., NOVA, R., BAÙ, A., and HAYKAL, R., 2007 – Prove di carico su pali di grande diametro mediante l’impiego di celle Osterberg, Rivista Italiana di Geotecnica (RIG), Anno 41, no. 4, pp 9-28, in Italian BOLTON, M. D., 1986 – The strength and dilatancy of sands, Géotechnique, vol. 36, No. 1, pp 65-78. CEN, 2004 - EN 1997-1:2004: Eurocode 7: Geotechnical design – Part 1: General rules, Brussels. CERNAK, B., 1976, The Time Effect Suspension on the Behavior of Piers, Institute of Civil Engineering, Bratislava, CSSR, Proc. 6th European Conference on Soil Mechanics and Foundation Engineering, Vienna, Vol. 1, pp. 111-114. EVANGELISTA L., PETRANGELI M. P., TRAINI G., 2003 –The cable-stayed bridge over the PO river, IABSE Symposium on structures for high-speed railway transportation, Antwerp, August, pp 138-139. FELLENIUS, B. H., ALTAEE, A., KULESZA, R, and HAYES, J., 1999 – O-Cell Testing and FE

analysis of a 28 m Deep Barrette in Manila, Philippines, ASCE Journal of Geotechnical and Geoenvironmental Engineering, Vol. 125, No. 7., pp 566-575. FLEMING, W. K. and SLIWINSKI Z. J., 1977 – The Use and influence of bentonite in bored pile construction, – DOE / CIRIA Piling Development Group Report PG 3. NG, C. W. W. and LEI, G. H., 2003 – Performance of Long Rectangular Barrettes in Granitic Saprolites, ASCE J. Geotech. and Geoenvir. Engrg., Volume 129, No. 8, pp. 685-696 O’NEILL, M. W. and REESE, L. C., 1999 – Drilled Shafts: Construction Procedures and Design Methods, report no. FHWA-IF-99-05, U.S. Department of Transportation, Federal Highway Administration. OSTERBERG, J. O., 1989 – New Device for Load Testing Driven Piles and Drilled Shafts Separates Friction and End Bearing, Proc. International Conference on Piling and Deep Foundations, London, A.A. Balkema, pp 421–427. RANDOLPH, M. F., 2003 – Science and empiricism in foundation pile design. Géotechnique, vol. 53, No. 10, pp 847–875.