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EXPERIMENTAL AND NUMERICAL ASSESSMENT OF OSTERBERG LOAD TESTSON LARGE BORED PILES IN SAND

Roberto Nova, Politecnico di Milano, ITALYBruno Becci, Ce.A.S. S.r.l., Milano, ITALY

The selection of the Osterberg Cell (O-Cell) technology as the standard testingmethod for deep foundations of the Railways Po Viaduct in Northern Italy offereda relevant number of field measures on large bored pile behaviour in alluvialsoils. The assessment of pile response during such non standard testingprocedure was performed through the comparison with standard top load testson companion piles as well as by numerical models of the tests including asimplified pile-soil interaction scheme. As shown in this paper, the O-Celltechnology allowed a careful assessment of the non-linear pile behaviour even atquite small loading levels as those required by a posteriori proof tests onproduction piles. Simple numerical models proved to be very effective in thesimulation and the interpretation of pile behaviour during such unconventional

testing procedure.

1. INTRODUCTION

In the construction of the High Speed RailwayViaduct crossing the Po river near Piacenza(Italy), large diameter bored piles were routinelyadopted as foundation system.

High design loads and complex site conditionssuggested the use of the Osterberg Cell (O-Cell)technology (Osterberg (1989)) as the first choicetesting method for these piles. At onshoreviaduct piers, usual kentledge load tests were

also performed, thus permitting worthwhilecomparison between different testing methods.

All the performed tests were also reproducedthrough simple numerical models that offered auseful and thorough assessment of themeasured pile behaviour during the O-Cell testsand a confirmation of the post-processingprocedures employed in their interpretation.

In the following, a review of the pile designcriteria is reported; then a discussion of bothdesign and proof load tests is included;numerical analyses of all the performed tests are

presented and relevant parameters that best fitexperimental observations are outlined. In thelight of all these observations, some generalremarks on the design of large diameter piles insand are proposed.

2. SITE AND VIADUCT DESCRIPTION

The Po Viaduct (Evangelista et al. (2003))includes 23 bays, of which 20 approaching theriver at both sides. A cable-stayed bridge, whosecentral bay is 192 m long, is placed at thepermanent riverbed crossing (fig. 1).

Figure 1: the Po Viaduct near Piacenza (Italy)

Two meter diameter bored piles were used atthe base of all the 24 piers, with pile lengthsranging from 40 to 62.5 m to ensure allowableloads between 10 to 18 MN. Bentonite slurry

was employed in borings. Pile number per piervaries from 4 to 28.

Nova. R, Becci B., Experimental and Numerical Assessment of Osterberg Load Tests on Large Bored Piles in

Sand, 33rdAnnual 11thInternational Deep Foundations Institute Conference Proceedings, New York, NY,

Oct 15-17 2008, pp 225-233.

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Subsoil conditions at the site are represented byvery deep submerged alluvial deposits: currentlya shallow 15 to 20 m thick sand layer isoverlying 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 mainlyinvestigated by means of SPT and CPT teststhat revealed almost normally consolidatedsands; limited to offshore piers, very deep layerswere sampled through a special CPT test, usingself-penetrating equipment into advancing hole.

In the granular deposits, SPT blow count Ncould be approximated by the followinganalytical expression:

N=(0.78 to 1.0) z (1)

in which z is the depth in [m]; typical relativedensities between 50% and 60% were estimatedand an almost constant peak friction angle equalto 31 was evaluated, including dilatancyreduction due to high pressure, according toBolton (1986).

The intermediate, slightly overconsolidated, siltyclay layer was analyzed through both in-situCPT and lab tests, showing characteristicundrained shear strength values ranging from 50to 100 kPa at depths between 15 and 30 m.

This intermediate cohesive layer plays animportant role in the hydraulic design of theoffshore piers, as it limits the expected scourdepth during extreme design flood conditions.

3. PILE DESIGN CRITERIA AND PRELIM-INARY EXPERIMENTAL FINDINGS

Preliminary design criteria are discussed inBecci et al. (2007). In Table 1 (left column) asummary of initial design assumptions isincluded.

Unit shaft resistance in sands was assumedlinearly varying with depth through a constantcoefficient K set to 0.6 and a friction parametertan(): as for , the critical friction angle of soil,set to 30, was selected. Such unit resistance

was assumed to develop at a differential pile-soillocal movement of 0.5% of pile diameter D. Asfor toe resistance, the design value in Table 1was considered to develop corresponding with atoe displacement of 5% D.

Before actual production pile constructionstages, two design load tests were performed,according to the ASTM D1143 Quick Load TestMethod, on sacrificial piles (50 and 55 m long),near actual Viaduct pier on the left handriverside, i.e. at the north side of Po river. Bymeans of a pair of Osterberg cells, installed ineach shaft, as shown in fig. 2, ultimate loads

could be almost reached.

-21 m

O-cells

SAND

SILTY CLAY

SILTY SAND

-57 m

-52 m

-36 m

-55 m

-45 m

-50 m

-42 m

-2 m0 m

D=2 m

PILE A PILE B

Figure 2: preliminary pile load test assembly

In both piles, the lower O-Cell was placed 2 mabove the toe, thus allowing a close depiction oftoe behaviour. Moreover, through the installationof ten strain gauge levels along the shaft, sheardistribution at relevant test stages was obtained

Table 1

Preliminary Design assumptions Preliminary load tests results

Shaftresistance

Sands qs= Ktan()v= 0.346v (2)Clay qs0 (3)

Sands qs= (0.62 to 0.85) v (5)

Clay qs(0.20 to 0.25) v (6)

Toeresistance Sands qbNqv10v (4) Sands qb(7 to 9) v (7)

v =effective overburden stress = z = buoyancy unit weight= 10 kN/m3

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by the measurement of the pile axial straindistribution.

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 toeresistance at 5%D displacement wasmeasured, whilst a higher shaft strength couldbe observed, the latter, however obtainedthrough unusually large shaft displacements thatcould be imposed thanks to the particular testingmethod offered by the O-Cell technology.Increasing shaft resistance with depth wasconfirmed.

The inspection of the equivalent top load curvesobtained by an a posteriori processing of

measures (solid gray curves in fig. 3) shows thatpreliminary design assumptions (dashed lines)had predicted a lower ultimate loading due to anunderestimate of shaft resistance.

0

10

20

30

40

50

60

70

0 50 100 150 200 250

PILE HEAD SETTLEMENT- [mm]

TOPLOAD

[MN]

-----

PRELIMINARY DESIGN CURVES

NUMERICAL BACKANALYSIS

O-CELL EQUIVALENT CURVES(POSTPOCESSING FROM TESTRESULTS)

PILE A

PILE BPILE A

PILE B

Figure 3: Top Load vs Top Displacementcurve for preliminary test piles

On the other hand, it was noted that thepresumed pile stiffness at working load level hadbeen well matched (see fig. 3, again); moreover,actual pile construction would have implied moredifficult working conditions than those occurred

during test pile installation; finally, long termextreme scour conditions for offshore pierswould have represented an important issue thatcould have been hardly investigated by similarpreliminary tests.

For all these reasons, in spite of the apparentlyconservative preliminary design criteria, no pile

length was reduced in the final design. Thischoice demonstrated to be wise, in the light ofactual construction observations.

4. NUMERICAL SIMULATION OF PREL-

IMINARY LOAD TESTS

In the practice, the simulation of the soil-pileinteraction by means of non-linear Winklersprings (ONeill & Reese, (1999)) is still probablythe most popular method to analyse single pilebehaviour in inhomogeneous soil conditions.

Adopting this approach within a finite elementframework, various pile and loading conditionscan be easily modelled, including the simulationof an Osterberg test as well.

Using this method, the dark solid curves in fig. 3have been computed, including strength andstiffness parameters summarized in fig. 4.

The toe reaction pattern has been included byassigning the behaviour as directly measured byO-Cells tests.

At 5%D toe displacement, the ratio Nq=qb/vwas found to be about 8.6 for Pile A and 6.7 forPile B.

SILTYCLAY

SAND10

20

30

40

50

60

10

20

30

40

50

60

200 200400 400q [kPa]s q [kPa]s-2 m

-52 m

-57 m

-21 m

-36 m

0 m-2 m

403

30060

40

130170

60

40

180

400

=3

=1

=4 =2.5

=1.5

=1

0.4

0.8

1.2

1.6

0.

SHAFT s / (D)

0.5% 1% 1.5% 2%

qs

q/

CLAY

SANDS

5% 10% 15% 20%

2

0

4

6

8

Basereactionq

[MPa]

b

PILE B

PILE A

BASE s / Db

PILE A PILE B

measured

measured

SILTYSAND

assumed innumericalanalysis

assumed innumericalanalysis

Figure 4: numerical back-analysis on preliminary

test piles

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As for shaft reaction, trial shear strength profiles(solid lines in fig. 4) along with a threshold sidedisplacement have been assigned. The latter isincluded via a scale factor , which is actuallyone of the free parameters in this back-analysis

process: note that, taking =1, the normalizedshaft displacement s/D as usually considered byONeill & Reese, (1999) is recovered.

Final set of data, giving a good reproduction oftarget results, shows that is currently higherthan usual values; as for pile B, shaft reactionwas found to be weaker yet stiffer than in pile A,in the deep sand layer. Improved numericalsimulations can be obtained provided a slightlyless smooth curve for shaft reaction in sand isadopted: this modification is not relevant forcurrent discussion, however.

5. PROOF LOAD TESTS WITH OSTER-BERG CELLS

In preliminary load tests, due to high expectedultimate loadings, the O-Cell technology wasdeemed almost mandatory; however, proof loadtests, up to 1.20 the maximum service pile loadNmax, could have been performed by means oftraditional methods as well, unless quitecomplex conditions would have to be dealt within offshore piers no. 7 and 8 (fig. 5).

Following some debate, three production piles inoffshore piers and two onshore piles (fig. 6)were equipped with one O-Cell only, properlydesigned to impose a sufficient equivalent loadfor proof purpose. Near to the onshore O-Cell

piles, three traditional top load tests were alsorequested.

Figure 5: site conditions for Piers 7 and 8

The O-cell was placed at about 20% of the pilelength above the toe, as deep as possible toavoid significant lateral pileresistancereductionand, in the same time, to allow desiredequivalent top load imposition without excessivecell loadings or toe movements.

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

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Dealing with one O-Cell only offered theopportunity to perform a numerical analysis ofthe test itself, with a straightforward modificationof the finite element model used to analyze thebehaviour of the pile subjected to vertical top

loads.To model the single O-Cell test, the finiteelement corresponding with cell position must beannealed and two equal loads must be appliedin opposite directions to the nodescorresponding with annealed element.

It should be noted that the simulation of a singleO-Cell test requires no complex constitutive lawfor soil springs, since all such elements basicallyundergo monotonic loading path only.

6. PROOF LOAD TESTS DISCUSSION

In the following, the results from each O-Cell testare discussed. Also these tests have beenconducted according to the Quick Load TestMethod.

Limited to tests 1 to 3 (listed in fig. 6),remarkable nonlinear aspects have beenhighlighted during the tests; therefore someconsiderations about ultimate loading can beattempted as well.

In onshore tests 4 and 5, nonlinear behaviour isless clear: on the other hand, a worthwhilecomparison with traditional load tests is

available, which can be considered acontribution to experimental assessment of O-Cell technology.

OFFSHORE PIERS

Fig. 7 outlines the results obtained for offshorepiles in piers 7 and 8, by plotting the O-Cell plateabsolute movements in upward (top plate) anddownward (bottom plate) direction.

In spite of applied (equivalent) loads slightlyhigher than maximum expected working loads,quite evident non-linear behaviour is obtained.

Whereas such a behavior would, in general, beundesirable in a proof test, it is currentlyacceptable in such a procedure: in particular, anumerical simulation of each of these tests cansufficiently explain that the observed behaviouris physiological.

For back-analyses, the pile-soil interactiondescribed in the previous section is used, withthe normalized curves shown in fig. 8, same tothose included in fig. 4, for shaft reaction.

In fig. 8 again, qsprofile and relevant parametersobtained through the back-analyses aresummarized; obtained responses are included infig. 7, superimposed to experimental data.

The back-calculated critical toe pressure qbLIM=4MPa, corresponding with a toe settlement equalto 0.05 D is a good assumption for all theanalyzed cases, and well agrees with eq. 7.

-50

-40

-30

-20

-10

0

10

20

30

0 5 10 15 20

O-Cell Gross Load (MN)

O-CellDisplacement[mm]

TEST1 - PIER 7 PILE 9

TEST 3 - PIER 8 PILE 20

TEST 3 - numerical

DOWN

UP

-50

-40

-30

-20

-10

0

10

20

30

0 5 10 15 20

O-C

ellDisplacement[mm]

TEST 2 - PIER 7 PILE 21

TEST2 - numerical

DOWN

UP

Figure 7 measured displacements and backanalysis results for offshore tests

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PILE HEAD +37.0 m

SILTY

CLAY

z [m]10

20

30

40

50

60

0 50 100 150 200

q [kPa]s

SAND

SILTY

SAND

O-CELL POSITION

TEST1&3

TEST2

1 1

1 1

0.751

1.5 1.5

TEST 1

TEST 2

q [MPa]b LIM

4 4

0.4

0.8

1.2

1.6

0.

SHAFT s / (D)

0.5% 1% 1.5% 2%

qs

q/

CLAY

SANDS

0.4

0.8

1.2

1.6

0.5% 10% 15% 20%

qbLIM

qb/

BASE s / Db

TEST 3

In TEST 1 or 3, the shaft resistances qsand theirmobilization levels quite well adhere to originaldesign assumptions (as included in Table 1, leftcolumn), except that a non-zero shaft resistanceis 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 thelower part, some reduced shaft resistances hadto be included, to match the observedbehaviour. At this stage, it is however hard tostate whether this diminished response is due tolocally weaker shaft resistance or to higher local

compliance. To reliably depict the actualbehaviour, the load would have had to beincreased to a much higher level.

The scaling parameter decreases with qs: thisnumerical effect is necessary to reproducesimilar side skin friction stiffness for all cases,independently from maximum strength value.

Anyhow, the back-figured parameters in fig. 8currently fall within typical ranges. It should benoted that the shear strength increase withdepth is confirmed: however, the mobilizationlevels tend to increase with depth as well, as

also noted in the analysis of preliminary loadtests discussed previously.

ONSHORE PIERS

In TEST 4 (fig. 9), both soil conditions and pilelength are very similar to those considered foroffshore piles.

The measured behaviour was consistently verysimilar to TEST 1 or TEST3: in particular theTEST3 back-analysis can reasonably reproducethe behaviour of this pile too.

Further modifications may be included, to bettermatch higher stiffness of the upper part as wellas a slightly lower stiffness of the lowersegment: however such changes would notmodify the overall description for this pile.

Finally, experimental results and numericalback-analysis of TEST 5 on the shortest pile inthis campaign, is reported in fig. 10, whilst back-figured parameters are included in fig. 11.

Back-figured shaft reaction displays a betterresponse than in offshore piles: in sand layers,qs from eq. (1) can be increased by a factor1.33, thus obtaining an average ratio=qs/v=0.46.

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

-30

-20

-10

0

10

20

30

0 5 10 15 20

O-Cell Gross Load (MN)

O-CellDisplacement[mm]

TEST 4 - PIER 9 - PILE 4

TEST3 - numerical

DOWN

UP

Figure 9 measured displacements andback analysis results for onshore TEST 4

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At adjacent pier no. 13 location (about 50 m farfrom pier 12, in S-E direction), corresponding

with same piling layout and very similar soilproperties, a traditional top load test wasperformed, using a steel kentledge.

These results could be effectively compared withboth equivalent top load curve obtained throughthe standard O-Cell post-processing procedureand numerical prediction based on back-figuredparameters from TEST 5, in fig. 11.

The comparison included in fig. 12 shows anexcellent agreement among experimental datafrom traditional test (dots), equivalent O-Cell

curve (solid gray line) and numerical prediction(solid black line).

The top load test could not investigate non-linearpile behaviour. Therefore, unless a top loading

could be revised as a more natural way to testactual pile behaviour, the information offered bythe traditional procedure is poor indeed ascompared with an O-Cell test.

The remaining two top load tests wereconducted at piers 11 and 18, on 46 m longpiles. In both tests, an almost linear behaviourcould be obtained, with top settlements of about3 mm for both piles, at a proof load of 12 MN.

7. DISCUSSION AND CONCLUSIONS

The selection of the Osterberg cell technologyallowed the conduction of load tests to very highload levels that would have been hardly imposeddue to complex environmental conditions.

The O-Cell method in proof load tests couldprovide engineers with more useful and preciseinformation than those currently available bytraditional methods. In particular, some non-

linear behaviour of part of the pile could beactivated even at quite low loads.

In addition to the routine post-processing ofsampled data as a part of the standard O-Cellprocedure, the authors performed simplenumerical simulations of the tests using theWinkler method, and found that observed

FREE FIELD +45.5 m

PILE HEAD +41.0 m

SILTY

CLAY

10

20

30

40

50

0 50 100 150 200q [kPa]s

SAND

SILTYSAND

O-CELL POSITION

TEST5

1 0.5

1

1.5

TEST 5

q [MPa]b LIM

4

z [m]

250

0.5

Figure 11: back analysis proof load testassumptions and results for TEST5

-20

-10

0

10

20

0 5 10 15 20

O-Cell Gross Load (MN)

O-CellDisplacement[m

m]-----

TEST 5 - PIER 12 - PILE 1

TEST 5 - numerical

DOWN

UP

Figure 10 measured displacements andback analysis results for onshore TEST 5

PROOF LOAD

11.91MN

0

2

4

6

8

10

12

14

16

0 3 6 9 12 15PILE HEAD SETTLEMENT [mm]

TOPLOAD[MN]

PRELIMINARYDESIGN ASSUMPTIONS

TOP LOAD TEST - PIER No. 13, PILE 1

TEST 5 - EQUIV. TOP LOAD CURVE BYO-CELL PROCEDURE

NUMERICAL ANALYSIS (TEST 5)

Figure 12 comparison between traditionaltop load test, O-Cell test and numericalprediction for onshore pile

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behaviour during these unconventional testscould be reasonably reproduced by means ofinteraction curves similar to those that wouldhave been used in a traditional pile model.

It is the authors opinion that these conclusionsmay contribute to increase the confidence bypracticing engineers in the selection of moderntesting techniques like the one discussed in thispaper.

It is important to realize that new testingmethods currently allow the conduction of pileload tests, in almost all the real conditions.Complex site conditions and/or very high loadscan, therefore, hardly be used as an excuse tolimit or even omit load tests at all.

On the other hand, a careful assessment of the

obtained results is always recommended. Inparticular, heavy modifications to initial design,based on reasonable and well-establishedassumptions, should be considered with greatcare.

As for the observed behaviour of these largediameter shafts, drilled under bentonite in sands,the authors found almost uniform toe behaviour,in good agreement with most frequently usedcorrelations in the practice.

As for shaft resistance, however, relevantdiscrepancies among design correlations,

preliminary load tests and final proof load testsfindings have been highlighted and discussed.

These findings should, in general, suggest aparticular care in the selection of shaftresistance parameters for bored piles in sand, allthe more because similar observations havebeen reported by others, regarding bored shaftsor barrettes in different soil conditions (e.g.Randolph (2003), Fellenius et al. (1999)).

Such discrepancies are primarily related toconstruction issues, which can be hardlyincorporated in preliminary design models (seeCernak (1976), Fleming & Sliwinski (1977), Ng &Lei (2003)).

Even in the light of these limited observations, itcan be argued that the suggested partial safetyfactors used in the growing LRFD method alsoin geotechnical engineering, may need somefurther discussion before being used in thepractice.

In particular, reference is made to Eurocode 7(CEN (2003)), that recommends a partial safetyfactor B=1.60 for toe resistance, higher than theshaft resistance factor S=1.30, for bored piles.

The aforementioned values were likely tuned toimplicitly account for different settlementsnecessary to activate each of the twocontributions. However the authors arguewhether such values may or may not conflictwith some actual findings like those reported inthis paper, as well as other frequent fieldobservations in practical pile constructions.

ACKNOWLEDGEMENT

The technical advice of Grandi Lavori Fincositstaff, leaded by dr. Augusto Ba and dr. RajiHaykal, is particularly acknowledged. Piling

Contractors TREVI S.p.A. and VIPP S.p.A.,General Contractor SNAMPROGETTI and theClient Italferr S.p.A. are also acknowledged, aswell as Loadtest Inc. engineers who providedand supported the Osterberg cell technology.

REFERENCES

BECCI, B., NOVA, R., BA, A., and HAYKAL,R., 2007 Prove di carico su pali di grandediametro mediante limpiego di celle Osterberg,Rivista Italiana di Geotecnica (RIG), Anno 41,no. 4, pp 9-28, in Italian

BOLTON, M. D., 1986 The strength anddilatancy of sands, Gotechnique, 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 Suspensionon the Behavior of Piers, Institute of CivilEngineering, Bratislava, CSSR, Proc. 6thEuropean Conference on Soil Mechanics andFoundation Engineering, Vienna, Vol. 1, pp.111-114.

EVANGELISTA L., PETRANGELI M. P., TRAINIG., 2003 The cable-stayed bridge over the POriver, 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

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analysis of a 28 m Deep Barrette in Manila,Philippines, ASCE Journal of Geotechnical andGeoenvironmental Engineering, Vol. 125, No. 7.,pp 566-575.

FLEMING, W. K. and SLIWINSKI Z. J., 1977 The Use and influence of bentonite in bored pileconstruction, DOE / CIRIA Piling DevelopmentGroup Report PG 3.

NG, C. W. W. and LEI, G. H., 2003 Performance of Long Rectangular Barrettes inGranitic Saprolites, ASCE J. Geotech. andGeoenvir. Engrg., Volume 129, No. 8, pp. 685-696

ONEILL, M. W. and REESE, L. C., 1999 Drilled Shafts: Construction Procedures andDesign Methods, report no. FHWA-IF-99-05,

U.S. Department of Transportation, FederalHighway Administration.

OSTERBERG, J. O., 1989 New Device forLoad Testing Driven Piles and Drilled ShaftsSeparates Friction and End Bearing, Proc.International Conference on Piling and DeepFoundations, London, A.A. Balkema, pp 421427.

RANDOLPH, M. F., 2003 Science andempiricism in foundation pile design.Gotechnique, vol. 53, No. 10, pp 847875.

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