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PROEFSCHRIFT

ter verkrijging van de graad van doctor aan de Technische Universiteit Delft,

op gezag van de Rector Magnificus, prof.dr.ir. J.T. Fokkema, voorzitter van het College voor Promoties,

in het openbaar te verdedigen op dinsdag 15 december 2009 om 15:00 uur door Jelke DIJKSTRA

civiel ingenieur geboren te Zwolle

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Prof.ir. A.F. van Tol

Copromotor: Dr.ir. W. Broere

Samenstelling promotie-commissie:

Rector Magnificus, Technische Universiteit Delft, voorzitter Prof.ir. A.F. van Tol, Technische Universiteit Delft, promotor Dr.ir. W. Broere, Technische Universiteit Delft, copromotor Prof.dr.ir. F. Molenkamp, Technische Universiteit Delft

Prof.dr.ir. W.S.J. Uijttewaal, Technische Universiteit Delft Prof.dr. D.J. White, The University of Western Australia Prof.dr. D. Muir Wood, University of Bristol

Dr.ir. O.M. Heeres, Gemeentewerken Rotterdam

Printed by:

W¨ohrmann Print Service P.O. Box 92 2702 CZ Zutphen The Netherlands Telephone: +31 575 58 53 00 E-mail: wps@wps.nl ISBN 978-90-8570-432-4 c ° 2009 by J. Dijkstra

All rights reserved. No part of the material protected by this copyright notice may be re-produced or utilized in any form or by any means, electronic or mechanical, including pho-tocopying, recording or by any information storage and retrieval system, without written consent of the publisher.

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Soft soil conditions require deep foundations for structures and buildings. The location of the bearing stratum, a soil layer that is capable of mobilizing enough bearing capacity to carry the load of the superstructure, is found at typical depths of 10 – 25 m in the western part of The Netherlands. The load from the superstructure, and in some cases the weight from the soil above the bearing stratum, is transferred to the bearing layer by piles.

The bearing capacity of a pile foundation is governed by the initial soil properties and the soil state (stress state and density) around the pile after installation. During installation displacement piles alter this soil state. However, this change in governing soil properties and the influenced area around the pile where this state is altered are still not well known.

An improved understanding of the soil behaviour during pile installation will probably lead to better predictions using finite element methods for pile bearing capacity of com-plex structures. This in turn may lead to optimizations of constructions involving large pile groups, i.e. quay-walls or high rise buildings, reducing the costs involved. Additionally, not only the prediction of bearing capacity of axially loaded compression piles could benefit from this research. A properly implemented installation phase will also improve the predic-tion of the capacity of other loading types, e.g. tension piles or horizontally loaded piles. The complementary situation where pile foundations are influenced by other construction activities can also benefit from a more realistic modelling of pile behaviour in finite element methods.

This thesis introduces two physical model tests to investigate the stress and density change in the soil during installation of a jacked pile as well as two numerical modelling approaches to model the same process in a finite element code capable of large deforma-tions.

In the first physical model test the stresses in the soil are measured by photoelasticity with a newly built full-field polariscope. The strains are calculated from the measured dis-placement fields, which are acquired with a digital image correlation (DIC) technique. The employed measurement methods limit the model setup to a plane strain configuration and a material substitution. The setup proved to be capable of measuring the stress distribution and strain distribution in the soil near the pile. The stress fields measured with the pho-toelastic method are in good agreement with the separately measured boundary loads. The accuracy of the displacement fields measured in the same setup is somewhat hampered by the low contrast of the photoelastic material. Nonetheless, the strains calculated from the smoothed displacement fields show loosening below the pile base and densification in the side lobes, which is in good agreement with one test on a high contrast material. The strain results are consistent with the stress results, i.e. the location and size of the influenced soil

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are in good agreement.

In the second model test, executed in a geotechnical centrifuge, the soil density change near an advancing pile has been measured with an adapted resistivity method. The boundary conditions in the centrifuge test allow for a fully three dimensional failure mechanism to develop around the pile. However, in this setup the stresses in the soil cannot be measured at the same time.

Regardless of the loose or dense initial conditions, the soil near the pile shaft loosened significantly during monotonic pile jacking in the centrifuge tests. Distinct differences in porosity change between the different instrument levels, located at different distances from the pile base, are found in the beginning of the installation phase. These differences tend to become smaller with increased penetration. In the pile test, performed after the pile installation stage, the soil densified during the initial unloading. During loading in the remainder of the pile test no additional density change is registered. Only when the pile is subsequently unloaded, a density increase is observed again. For a series of displacement controlled loading–unloading cycles the largest densification is found for the first series of small amplitude cycles. The largest density changes results from the loading–unloading reversals.

The main observation is that an increase in soil density adjacent to the pile during cyclic loading is coupled to a decrease in shaft resistance. The stress build up near the pile shaft and the corresponding loosening of the soil during the pile installation stage is destructed by densification and loss of shaft resistance during the subsequent cyclic tests at large dis-placement amplitude (> 0.5 mm).

The jacked pile installation process has been successfully simulated by combining the density dependent strength and stiffness response of the hypoplastic constitutive model with a numerical framework capable of large deformations. Two modelling approaches have been investigated. The first approach simulates the pile installation process by inflow of material on the lower boundary while the pile is fixed. In the second approach the pile is pushed in the soil. In both approaches the mesh is fixed. The approach where the pile penetrates the soil produced a more continuous spatial distribution of the results.

The numerical simulations agree better with the centrifuge results than with the pho-toelastic tests. The calculated stress and density evolution for the centrifuge test is in reas-onable agreement with the measured response.

The results of the photoelastic model tests allow for a comparison of the simulated and measured spatial stress distribution. The calculated magnitude of the stress below the pile base is 3 to 5 times higher than measured. The low stress conditions in the test for the determination of the critical friction angle resulted in an overprediction of the friction angle and therefore the soil strength.

In the tests the parameters of the constitutive model are derived from laboratory tests and the initial conditions are obtained from the model tests. No additional changes are made in order to fit the experimental data. Therefore, both modelling approaches are applicable in practice with a comparable accuracy when the material properties and initial density are changed according to the in-situ conditions.

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In deltagebieden is de draagkracht van de bovenste grondlagen vaak onvoldoende om ge-bouwen direct op maaiveld te funderen. Daarom worden veelvuldig paalfunderingen toege-past, waarbij de belasting overgebracht wordt naar dieper gelegen, stijvere, lagen. In het westen van Nederland worden de bovenste draagkrachtige lagen aangetroffen op een diepte van 15 tot 25 meter onder maaiveld.

Of een bepaalde laag draagkrachtig genoeg is hangt af van de gebouwbelasting, het gebruikte type palen en de grondeigenschappen na paalinstallatie. De grondeigenschappen na paalinstallatie hangen af van de oorspronkelijke grondeigenschappen en de soil state (de grondtoestand: spannings situatie en de dichtheid) rondom de paal na installatie. Grond-verdringende palen veranderen deze soil state gedurende het installeren sterk. Echter, in welke mate de grondeigenschappen veranderen en in welk gebied rondom de paal deze eigenschappen worden be¨ınvloed is nog steeds onduidelijk.

Een verbeterd inzicht in het grondgedrag gedurende het installeren van grondverdrin-gende palen zal kunnen leiden tot betere voorspellingen met eindige elementen methoden voor complexe paal interactie problemen. Dat kan optimalisaties opleveren voor constructies met grote paalgroepen, bijvoorbeeld kademuren of hoogbouw funderingen, zodat de kos-ten worden teruggebracht. Naast een verbeterde voorspelling voor de axiale draagkracht van drukpalen kunnen andere belastingsgevallen ook profiteren van dit onderzoek. Een goed ge¨ımplementeerde paal installatie fase resulteert ook in verbeteringen van het gedrag van bijvoorbeeld trekpalen of horizontaal belaste palen. Het complementaire geval, waar een bestaande paalfundering wordt be¨ınvloed door andere bouwactiviteiten kan ook beter worden voorspeld, wanneer de paal installatie realistischer is gemodelleerd in een eindige elementen methode.

Om tot een dergelijke verbeterde modellering te komen, is een beter inzicht in de soil state tijdens en na paalinstallatie nodig. Daarnaast moet deze verandering van soil state meegenomen worden in een eindige elementen modellering.

Dit proefschrift introduceert twee fysische modelproeven die de spannings- en dich-theidsveranderingen in de grond rondom de paal gedurende installatie van een weggedrukte paal bestuderen.

In de eerste modelproef zijn de spanningen in de grond gemeten met een nieuw ont-worpen polariscope, welke het hele meetdomein rondom de paal kan bemonsteren, door gebruik te maken van foto-elasticiteit. In dezelfde opstellingen worden de rekken berekend uit, de met DIC (digital image correlation) gemeten, verplaatsingsvelden. Deze twee meet-methoden limiteren de modelopstelling tot een spleetmodel waar ook het grondmateriaal is vervangen door een foto-elastisch materiaal. Het is aangetoond dat met deze opstelling de spannings- en rekdistributie in de grond rondom the paal gemeten kan worden. De met

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de foto-elastische methode gemeten spanningsvelden komen goed overeen met de krachten op de rand, gemeten met een drukcel. De nauwkeurigheid van de in dezelfde opstelling gemeten verplaatsingsvelden wordt be¨ınvloed door het lage contrast van het foto-elastische materiaal. Desalniettemin laten de uit de gefilterde verplaatsingen berekende rekken een losser wordend pakket onder de punt zien en een verdichting in de zijflanken. Dit is in overeenstemming met de proef uitgevoerd op een hoog contrast grond materiaal. De rekken zijn consistent met de spanningen, oftewel de locatie en de grootte van het verstoorde gebied zijn vergelijkbaar.

In de tweede modelproef is de dichtheidsverandering rondom een penetrerende paal ge-meten met een resistiviteits meting. De resistiviteit wordt op drie hoogtes boven de paalpunt langs de paal gedurende het indringen gemeten. De proeven zijn uitgevoerd in een geotech-nische centrifuge. In deze proeven kan het bezwijkmechanisme alzijdig rondom de paal ontstaan. Echter, in deze opstelling kunnen de spanningen in de grond niet gelijktijdig worden gemeten. De dichtheid van de grond rondom de paalschacht reduceerde significant tijdens het wegdrukken van de paal in de centrifuge proeven, onafhankelijk van de initi¨ele losse of vast gepakte condities. De instrumentatie niveaus, gepositioneerd op drie verschil-lende afstanden van de paalpunt, laten in het begin van de proef duidelijke verschillen zien. Deze verschillen worden kleiner naarmate de paal dieper wordt weggedrukt. Gedurende de paalproef, die is uitgevoerd na de paalinstallatie, verdichtte de grond tijdens het ontlasten. Na deze ontlasting veranderde de dichtheid nauwelijks meer tijdens het herbelasten. Een verdichting is opnieuw waargenomen wanneer de paal opnieuw wordt ontlast. In een serie verplaatsingsgestuurde belasting–ontlasting cycli wordt de grootste dichtheidsverandering gevonden voor de eerste serie cycli met kleine amplitude. De grootste verandering in dich-theid is gevonden voor de wisselingen in belastingsrichting.

Het wegdrukken van palen is succesvol gesimuleerd wanneer het dichtheidsafhankelijk sterkte- en stijfheidsgedrag van het hypoplasticiteit model wordt gecombineerd met een voor een grote deformaties geschikt numeriek schema. Twee modelleringsmethodes zijn onderzocht. De eerste methode simuleert de paalindringing door rondom een gefixeerde paal materiaal aan de onderzijde van het domein in te stromen. In de tweede methode wordt de paal daadwerkelijk weggedrukt in de grond. Voor beide methode is de mesh stationair. De tweede methode, waarbij de paal in de grond gedrukt wordt, produceerde een betere ruimtelijke distributie van de resultaten.

De numerieke resultaten komen beter overeen met de gemeten resultaten in de centrifuge proeven dan de foto-elastische proeven. De berekende spannings- en dichtheidsontwikkel-ing in de gesimuleerde centrifugeproef komt redelijk overeen met de gemeten waarden.

De foto-elastische proefresultaten maken het vergelijken van de ruimtelijke distributie van de spanningen mogelijk. De berekende grootte voor de spanningen onder de paalpunt ligt 3 tot 5 keer hoger dan is gemeten. De lage spanningen gedurende het bepalen van de kritische hoek van inwendige wrijving heeft geleid tot een overschatting van de wrijving-shoek en daarmee de sterkte van de grond.

De eigenschappen van het constitutieve model zijn bepaald uit laboratoriumproeven en de initi¨ele condities zijn bepaald uit de modelproeven. Deze parameters zijn niet ver-anderd om de experimentele resultaten beter te benaderen. Beide modelleringen zijn daarom toepasbaar in de praktijk met een vergelijkbare nauwkeurigheid wanneer de materiaalei-genschappen en initi¨ele dichtheid worden aangepast aan de in-situ eimateriaalei-genschappen.

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List of Symbols xiii

1 Introduction 1

1.1 Background . . . 1

1.2 Objectives & Limitations . . . 3

1.3 Outline . . . 4

2 Pile Installation 5 2.1 Introduction . . . 5

2.2 Prediction Methods for Pile Bearing Capacity . . . 6

2.2.1 Limit State: Direct Methods . . . 6

2.2.2 Limit State: Indirect Methods . . . 7

2.2.3 Finite Element Analysis . . . 8

2.3 Effects of Pile Installation on Bearing Capacity and Stiffness . . . 10

2.3.1 Bearing Capacity . . . 10

2.3.2 Pile Stiffness . . . 12

2.4 Stress Change . . . 13

2.4.1 Local shaft friction . . . 13

2.4.2 Horizontal Contact Stresses . . . 15

2.4.3 Overview . . . 18 2.5 Density Change . . . 19 2.5.1 Three Dimensional . . . 19 2.5.2 Plane Strain . . . 21 2.5.3 Overview . . . 23 2.6 Soil State . . . 25

2.7 Long Term Effects . . . 27

2.8 General Conclusions . . . 29

3 Measurement of Stress and Density Change 31 3.1 Introduction . . . 31 3.2 Principles of Photoelasticity . . . 31 3.2.1 Theory . . . 32 3.2.2 Measuring Photoelasticity . . . 38 3.2.3 Phase Unwrapping . . . 43 ix

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3.2.4 Stress Separation . . . 47

3.2.5 Testing of Phase Unwrapping and Stress Separation Algorithms . . 49

3.3 Principles of Digital Image Correlation . . . 52

3.3.1 Image Acquisition and Post-Processing . . . 53

3.3.2 Peak Finding Algorithm . . . 54

3.3.3 Strain . . . 55

3.4 Measuring Density Change by an Electrical Analogue . . . 56

4 Photoelastic Investigation into Pile Installation 61 4.1 Introduction . . . 61 4.2 Model Scaling . . . 61 4.2.1 Material Substitution . . . 62 4.2.2 Geometrical Considerations . . . 62 4.2.3 Stress . . . 63 4.2.4 Conclusions . . . 65

4.3 Model test setup . . . 65

4.3.1 Mechanical Realization . . . 66

4.3.2 Control & Measurement Setup . . . 67

4.3.3 Preparation . . . 70

4.3.4 Workflow . . . 71

4.4 Test Series . . . 72

4.5 Results . . . 73

4.5.1 Evolution of Pile Head and Surcharge Load . . . 73

4.5.2 Full Field Stress Evolution . . . 76

4.5.3 Full Field Strain Evolution . . . 82

4.6 Discussion . . . 87

4.7 Conclusions . . . 87

5 Centrifuge Model Pile Tests 89 5.1 Introduction . . . 89 5.2 Model Scaling . . . 89 5.2.1 Stress . . . 90 5.2.2 Geometrical Considerations . . . 90 5.2.3 Groundwater . . . 91 5.2.4 Geotechnical Centrifuge . . . 91

5.3 Model Test Setup . . . 92

5.3.1 Mechanical Realization . . . 92

5.3.2 Measurement Setup . . . 94

5.3.3 Model & Sample Preparation . . . 97

5.3.4 Estimated Measurement Error . . . 97

5.4 Test Series . . . 98

5.5 Results . . . 99

5.5.1 Axial Forces during Pile Installation . . . 99

5.5.2 Axial Forces during Pile Load Tests . . . 100

5.5.3 Axial Forces during Cyclic Tests . . . 103

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5.5.6 Density change during the Pile Load Tests . . . 111

5.5.7 Densification during the Cyclic Tests . . . 112

5.6 Discussion . . . 115

5.7 Conclusions . . . 115

6 Numerical Simulation of Pile Installation 117 6.1 Introduction . . . 117

6.2 Large Deformations . . . 117

6.2.1 Geometrical Non-Linearity . . . 117

6.2.2 Fixed Mesh . . . 118

6.3 Constitutive Model: Hypoplasticity . . . 119

6.4 Groundwater Flow . . . 120

6.5 Modelling Pile Installation . . . 120

6.5.1 Fixed Pile . . . 121

6.5.2 Moving Pile . . . 121

6.6 Simulation of Photoelastic Tests . . . 122

6.7 Simulation of Centrifuge Tests . . . 123

6.8 Overview Calculations . . . 126 6.9 Results . . . 127 6.9.1 Photoelastic Tests . . . 127 6.9.2 Centrifuge Tests . . . 130 6.10 Discussion . . . 138 6.11 Conclusions . . . 138

7 Conclusions and Recommendations 141 7.1 Conclusions . . . 141

7.2 Recommendations . . . 143

Bibliography 144 A Phase Unwrapping Algorithm 161 A.1 Lp-norm Algorithm . . . 161

A.2 PCG Algorithm . . . 163

B Galvanometer Circuit 167 C Measurement Consistency 169 D Additional Photoelastic Test Results 171 D.1 Stresses . . . 171

D.2 Strains . . . 174

E Effect of Subwindow Size in DIC Analysis 177

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G Parameter Derivation 183

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Latin

a polarization vector

A displacement gradient tensor

B magnetic flux intensity

C stress-optical coefficient

Cst storage coefficient

C0 photoelastic constant

d50 mean grain size

D pile diameter

D electric displacement field

e void ratio

ed, ec, ei minimal, critical, and maximum void ratio

ed0, ec0, ei0 low stress limits of minimal, critical, and maximum void ratio

E Young’s modulus

E electric field strength or Biot strain tensor

F deformation gradient tensor

g gravitational acceleration

h hydraulic head

hs granulate hardness

H magnetic field strength

I light intensity

I identity matrix

Jf electric current

k hydraulic conductivity

k, l, m wave vectors

K lateral earth pressure coefficient

K intrinsic permeability

L pile length

n porosity or exponent of granulate hardness

N blow count for standard penetration test (SPT)

Nq bearing capacity factor

ptotal total groundwater pressure

P electric polarization tensor, Jones Matrix for a linear polarizer qb, qc, qs base, cone and shaft resistance

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Qdriven, Qjacked static bearing capacity for driven, jacked piles

R electical resistance

Rn normalized apparent resistivity

R Jones matrix for a linear retarder, rotation tensor

Ts Cauchy stress tensor

˚Ts objective stress rate tensor

u (pile) displacement

U stretch tensor

vm material velocity

vg groundwater velocity

Vs voltage potential in soil

Vc voltage potential in current source

w gradient weight

Greek

α, β exponents controlling volumetric behaviour in hypoplastic law χe electric susceptibility

δ retardation angle

δJ total variation of functionJ

∆ retardation of non-ideal retarder

∆x, ∆y wrapped phase difference in horizontal and vertical direction

ǫ0 permittivity of free space

ǫr dielectric tensor

ε i = j normal strain component, i 6= j shear strain component ηc complex refractive index tensor

η1, η2, η3 principal refractive indices

γ conductivity

κ unwrapped phase data

µ0 permeability of free space

µr relative permeability

ν Poisson ratio

φ isoclinic angle

φc critical friction angle

ρf free charge density

ρgw groundwater density

ρs soil resistivity

̺ charge boundary

ρf free charge density

ρgw groundwater density

ρs soil resistivity

̺J total variation of functionJ

σc,σres confining pressure, residual stress

σh, σv horizontal, vertical stress component, (’) denotes effective stress

σij i = j normal stress component, i 6= j shear stress component

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Introduction

1.1

Background

Soft soil conditions require deep foundations for structures and buildings. The location of the bearing stratum, a soil layer which is capable of mobilizing enough bearing capacity for the superstructure, is found at typical depths of 10 – 25 m in the western part of The Neth-erlands. The load from the superstructure on the surface, and in some cases the weight from the soil above the bearing stratum, is transferred to the bearing layer by piles. Traditionally, wooden piles were used. Nowadays most piles are made out of (pre-stressed) reinforced concrete or steel.

Over the years different pile types were developed. These pile types differ in material use, geometry, and in construction method. Prefabricated piles, for example, are installed into the soil after production in a factory, whereas other pile types e.g. the continuous flight auger pile (CFA) is cast in place. Most prefabricated piles are driven into the ground with an impact hammer. Each blow of the ram displaces the pile deeper into the soil until the target depth is reached.

If, alternatively, the pile is jacked, the pile is pushed into the soil with hydraulic jacks. The total pile length prohibits the installation of the pile in a single stroke, therefore the pile will be installed in several strokes of 0.5 – 1 m. Jacked piles are also called pressed-in piles. In contrast with driven piles a reaction force needs to be mobilized to prevent uplift of the equipment. A significant difference in vibrations, emitted from the equipment and pile during the pile installation process, exists between the driven and jacked piles. The jacked pile installation causes limited additional vibrations, i.e. the process mainly comprises of pile and soil displacements. The description of the pile driving process is complicated by the stress waves in the pile and soil resulting from the driving process. Piles installed either driven or jacked are displacement piles, as no soil material is removed during pile installation.

The bearing capacity of a pile foundation is governed by the soil properties. The pile response is influenced by the strength and stiffness of the soil, and also by the properties of the pile-soil interface. As non-cohesive soils are particulate materials, the continuum soil properties originate from the grain properties as well as the assembly, and the contact forces

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between the grains in the assembly of grains. Therefore, sand exhibits stress related soil properties. Higher stress state, or larger contact forces, yield higher strength and stiffness. This density and stress dependency of the strength and stiffness is sometimes referred to as soil state. Exactly this soil state is altered by the installation process of displacement piles.

The change in the governing soil properties, or soil state, and the influenced area around the pile where this state is altered are still not well known. The following chapter will show the difficulties and uncertainties involved in the characterization of the change in soil state resulting from and during pile installation.

In engineering practice these effects are accounted for in empirical design methods. However, most of these methods only estimate the bearing capacity and do not model the underlying physical mechanisms.

The existing empirical methods are sufficiently accurate for the common foundation designs in which piles are only loaded in axial compression and influences on existing build-ings are negligible. The initial stiffness response, the load settlement behaviour of the pile head, is already much harder to predict. For situations where interaction between founda-tion elements is important, e.g., a quay wall with both tension and compression piles and the laterally loaded wall itself, or new pile foundations near existing piles, these prediction methods are of very limited use. These situations occur regularly for pile groups or when piles are installed near existing buildings in an urban environment.

Ideally, the prediction of pile bearing capacity has to incorporate the changes in soil properties to properly predict the pile bearing capacity. When modelling pile foundations in a finite element code the installation phase is often not incorporated. The change in soil properties is not accounted for, nor is the full simulation of the installation phase itself. This oversimplification leads to large differences between finite element predictions of pile behaviour and measurements [39].

In order to improve the prediction of pile bearing capacity from finite element methods the altered soil conditions from the pile installation stage should be accounted for in the calculations. Two main problems complicate this:

- Firstly, there is a poor understanding of the soil behaviour near the pile during pile installation and after pile installation has finished. These complex effects should be studied, in order to develop or select the proper constitutive model for the soil and to validate the results from the numerical simulations;

- The second problem originates from the numerical implementation of common small strain finite element methods. The numerical framework of these methods cannot handle large (local) deformations that occur during pile installation. These large local deformations also influence the porewater pressures. Therefore the numerical code should also incorporate a coupled material and groundwater description. For pile driving the inertial effects in the pile and soil should also be addressed, which most numerical codes cannot deal with either.

An improved understanding of the soil behaviour during pile installation will lead to better predictions from finite element methods of pile bearing capacity for complex struc-tures. This in turn should lead to optimizations of constructions involving large pile groups, i.e., quay-walls or pile rafts, possibly reducing the costs involved. Additionally, not only

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research. A properly implemented installation phase will also improve the prediction of the capacity of other loading types of, e.g. tension piles or horizontally loaded piles. The com-plementary situation where pile foundations are influenced by other construction activities can also benefit from a more realistic modelling of pile behaviour in finite element methods. This thesis is part of a larger research project in which the implementation of the pile installation stage of jacked and driven displacement piles in finite element codes is invest-igated. Next to the theoretical and numerical implementation aspects, the soil behaviour during and after pile installation is also experimentally investigated.

1.2

Objectives & Limitations

The general objective of this study is to quantify the change in soil state around the pile as a result of pile installation. This study is limited to the behaviour of jacked displacement piles in sand. Dynamic effects of pile driving and long-term pile setup effects are not taken into account. Given the problems presented in the previous section, the following objectives are formulated:

- The changes of density and stress in the soil near the pile are investigated in physical experiments. The density response below the pile base and next to the pile shaft as well as the stress change in these zones will be studied in 1g plane strain model tests. The density change near the pile shaft will be investigated in geotechnical centrifuge tests.

- The physical model tests are simulated in a finite element code capable of large de-formations. Results of the experiments will be used to support the selection of a proper constitutive soil model for these simulations. The results of the simulations will be used for validation of the numerical modelling and the extension of the meas-urement results to in-situ conditions.

Physical model tests are selected, because the measurement data from instrumented field tests is limited by the in-situ test conditions, which are difficult to control. The accurate measurement of these conditions, especially the change in soil properties, is close to im-possible. Model pile tests, on the other hand, allow for control of the initial soil conditions and the measurement of the change in these soil properties.

Each model pile test has model scale limitations from improperly scaled stress state and soil properties. The benefits of more detailed measurement data sometimes justifies poor scaling. An example of such a model test is the photoelastic test setup presented in this research, where the sand is substituted by glass and the experiment is reduced to a plane strain setup. At the same time the technique allows for the detailed measurements of stress and density changes in the soil next to the pile. On the other hand, the geotechnical centrifuge tests, also part of this research, properly apply the scaling rules at the cost of detailed information on the stress state in the soil during these tests. These tests present the measured porosity change in the soil near an advancing displacement pile. Together these two physical model tests should provide an overview of the soil behaviour near a

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displacement pile. The numerical simulations will help in scaling the model results to the in-situ case, as well as providing additional support for the interpretation of the measurement data.

1.3

Outline

The thesis consists of 7 chapters. In Chapter 2 the phenomenon of pile installation is dis-cussed. In Chapter 3 the theoretical principles of the applied density and stress measuring methods are elaborated. Chapter 4 presents the experimental results of photoelastic model tests in which both density and stress change near a displacement pile are measured. Chapter 5 presents a series of geotechnical centrifuge tests in which the axial forces acting on the pile and the porosity change of the soil near the pile are measured. In Chapter 6 the find-ings of the experiments are compared with a series of numerical simulations of the pile installation process. Finally, Chapter 7 sums up the findings of this study and gives some recommendations for future research.

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Pile Installation

2.1

Introduction

Before a prefabricated displacement pile is functional, the pile is first installed into the ground. Almost all of the prefabricated piles are driven into the ground with an impact hammer. The kinetic energy from each blow of the ram displaces the pile deeper into the soil. A jacked pile is pushed into the soil with hydraulic jacks, again the pile displaces the soil. In contrast with driven piles a reaction force needs to be mobilized to prevent uplift of the equipment. A significant difference in vibrations radiated from the equipment and pile during the pile installation process exists between the driven and jacked piles during pile installation. The jacked pile installation does not emit much additional vibrations, i.e., the process mainly comprises of pile and soil displacements. Whereas, the description of the pile driving process is complicated by the stress waves in the pile and soil that results from the driving process. Piles installed with these installations methods are displacement piles; no soil material is removed during pile installation.

The ultimate bearing capacity of a pile foundation is mainly governed by the soil prop-erties. Not only the strength and stiffness response of the soil, also the pile-soil interface properties need consideration. The stiffness and strength properties are continuum descrip-tions of the soil behaviour. In a non-cohesive soil the continuum soil properties originate from the grain properties as well as the assembly and contact forces between the grains in the assembly of grains. Sand, therefore, exhibits stress related soil properties, i.e., an higher stress state, or larger contact forces, yields higher strength and or stiffness. Strength and stiffness depend on the density and stress of the soil; sometimes referred to as soil state. Exactly this soil state is altered by the pile installation process.

The change in magnitude of the governing soil properties or soil state and the influence area around the pile where this state is altered are still not well known. How do these effects influence the bearing capacity? Ideally, the evolution of the stress and density in the soil is monitored during the pile installation stage. Unfortunately, existing research studies only one aspect, i.e. stress change or density change.

This Chapter will present how current design methods incorporate the pile installation stage in the prediction of bearing capacity, if at all. Also, the magnitude of the change in

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density or in stress or in both is investigated by studying existing experimental results. The focus is on pile installation, i.e., the implementation of pile installation in design methods, the effect of pile installation on: bearing capacity, the stress change in the soil, the density change in the soil and the combined change in stress state and density change.

2.2

Prediction Methods for Pile Bearing Capacity

Three distinct approaches for the prediction of pile bearing capacity can be identified: (1) Direct limit state methods correlate in-situ tests directly to pile bearing capacity, e.g., the Koppejan method commonly used in the Netherlands. (2) Indirect limit state methods, e.g., limit equilibrium methods and cavity expansion solutions, first derive model parameters from in-situ or laboratory tests and subsequently use these in empirical and analytical meth-ods. (3) A finite element analysis uses the soil properties obtained from the laboratory tests and the implemented constitutive model. Of these approaches the first two are most of-ten used in practice. Still a lot of empiricism is found in foundation design [150]. This is reflected in the prevalence of the first two methods.

2.2.1

Limit State: Direct Methods

The first, developed design methods, are all direct limit state methods, in which the results from the in-situ measurement are directly converted to an predicted maximum allowable pile bearing capacity. In practice these methods have proven to be successful for single piles. Various in-situ measurement instruments are used as input for the design of pile foundations [33]; the cone penetration test (CPT), the standard penetration test (SPT), and the pressuremeter test (PMT). The last method, however, is not used as a direct correlation. In direct methods the in-situ installation effects, arising from the test, are directly accounted for in the prediction of pile bearing capacity. This is a major advantage compared with the indirect methods.

Many correlations between the measured cone resistanceqc or the blow count from

the standard penetration testN and pile bearing capacity have been developed in the past decades for different subsoils and different pile types. Internationally, the Dutch method [156], the French method [41], the method of Schmertmann [163] and the method proposed by Eslami & Fellenius [73] are most established. More recently new methods have been developed, especially for large diameter closed- and open-ended offshore piles in sand [49], [96], [109], [114]. Of these methods only the Dutch and French methods directly account for pile installation effects, whereas in [114] only the soil displacements of the displaced soil volume of the pipe pile are accounted for. However, the offshore design methods are specifically designed for driven piles, so they are tailored for one installation process.

An overview of CPT methods and their performance focussing on mechanical and elec-trical CPTs in mixed soil conditions for different pile lengths is given in [36] and [73]. The prediction method is using an effective cone resistance reading by accounting for pore wa-ter pressures and a geometric averaging (ageo = (a1a2. . . an)1/n) method instead of an

arithmetic averaging (aari = (a1+ a2. . . + an)/n) method. Geometric averaging is less

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The influence of pile installation is sometimes lumped together with other corrections in a single empirical reduction factor. In a recent evaluation of the performance of offshore pile design methods, [204], the Dutch averaging method of the cone resistance readingqcis

found to be more accurate for the large diameter piles. The method of [73] was not taken in the comparison.

Meyerhof [130] presented an empirical relation between SPT blow count and pile ca-pacity. The pile type and pile installation are accounted for in separate empirical reduction factor for the shaft and base resistance. These reduction factors depend on the type of in-stallation. For driven and bored piles different factors have been proposed.

2.2.2

Limit State: Indirect Methods

Indirect methods use the strength and stiffness properties derived from laboratory tests or derived model parameters from in-situ tests as input for calculations for the prediction of pile bearing capacity. Limit equilibrium methods as well as more advanced cavity expansion methods are used for this purpose.

Similar to shallow foundations, bearing capacity factors have been analytically derived from limit equilibrium analysis using classic plasticity theory and a pre-defined slip plane (e.g. [181], [22]). When combined with the vertical effective stress and the pile base area a static pile base capacity can be predicted. Some authors introduce more complex soil models and have derived the bearing capacity factor numerically [86]. For the shaft resistance of piles in sands an estimated effective horizontal stress in combination with the effective pile-soil friction angle is commonly used. The effective horizontal stress is derived from the initial effective vertical stress with the coefficient of lateral earth pressureK. From pile load tests is found thatK depends on pile type [124]. The following values are found: H-pilesK = 1.4−1.9, pipe piles K = 1.2−1.3, precast square concrete piles K = 1.45−1.6, timber piles (1 test)K = 1.25, tension piles (all types) K = 0.4 − 0.9. In this method the effects of pile installation are hidden in the empirically derived recommendations for K from pile load tests on several pile types. A full overview of limit equilibrium methods for all soil types is given in [33].

The cavity expansion solutions have been developed mainly to replace the rather simple pre-defined slip planes in limit equilibrium analysis by introducing a more realistic failure mode, e.g., cylindrical or spherical cavity expansion. The cavity limit pressure is correlated with cone resistanceqcand subsequently used to derive pile bearing capacity. The cavity

methods are often used for the interpretation of cone penetration tests, and to a lesser extent for the prediction of bearing capacity of pile foundations.

The method is originally developed for metal research, e.g., for a spherical expansion in a Tresca material [90]. In soil mechanics the method is introduced for the analysis of a spherical expansion in associated Mohr-Coulomb material [43]. The method has been extended with an approximate solution for a spherical expansion in a compressible Coulomb material with non-zero volumetric strains [186] or for an non-associated Mohr-Coulomb flow rule [42]. These solutions, however, are only valid for small deformations. It is of importance to cope with large deformations and volumetric changes of the soil. The large strain solutions for spherical and cylindrical cavity expansion are solved for

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non-cohesive soils and non-cohesive soils by solving for the non-associated Mohr-Coulomb flow rule [212] and a critical state model [51]. More recently ([160], [161]) the empirical dilatancy relationship of Bolton [31] is implemented in the spherical and cylindrical cavity expan-sion framework for a more accurate prediction of the limit pressures. In order to improve upon the volumetric soil behaviour a numerical solution of the large strain cylindrical cav-ity expansion problem, using a hypoplastic constitutive model which incorporates dilatancy, contractancy and the dependence of stiffness on stress and density, is developed in Germany [53], [55], [143].

In the past decades the methods have been extended from small to large strain and from rather crude elasto-plastic models to far more advanced constitutive models as the hypoplasticity model. The results are still of limited value as these methods do not properly simulate the complex soil behaviour near the pile shoulder or the tip of the CPT or the pile. Some methods indirectly relate soil properties to pile bearing capacity, by correlating an indirectly derived relative density from CPT readings [102] to pile bearing capacity [121]. In this method pile installation effects are not explicitly accounted for.

2.2.3

Finite Element Analysis

As opposed to direct limit state methods, that use correlations with in-situ tests to predict a maximum bearing capacity, and indirect limit state methods, that use model parameters derived from laboratory tests and limit state solutions from equilibrium analysis to predict maximum pile bearing capacity, finite element analysis (FEA) is not restricted for studying an ultimate bearing capacity. The results from FEA also show the calculated stress and strain distribution in the soil and in the pile and calculate the stiffness response of the pile.

In FEA the constitutive model models the material behaviour. This constitutive model requires model parameters for the soil and pile material and the interaction between soil and pile material. These properties are generally derived from laboratory tests performed on samples taken from the construction site before the piles are installed. The change in soil properties, therefore, needs to be simulated as well; these are not reflected in the material properties. The latter is automatically incorporated in the direct limit state methods.

In current practice the pile installation phase is often omitted when displacement piles are modelled in a finite element code. The pile is modelled at the installation depth by a beam element or isoparametric elements connected to the isoparametric elements of the soil. The model parameters of the soil are not corrected for pile installation.

The full simulation of pile installation is even more difficult to model. This is partly due to the inability of the numerical scheme used in today’s geotechnical finite element applications and partly due to the complex soil behaviour around a penetrating displacement pile. The total Lagrangian and updated Lagrangian description are most commonly used in geotechnical finite element applications. The weakness in both methods is the inability to track large deformations in the soil body. The numerical integration scheme for acquiring the displacements is only accurate or even stable for a limited distortion of the elements.

Some authors published ways to incorporate the effects of pile installation in a small strain code, [14], [140]. The installation effects are simulated by loading the boundary of the mesh and applying additional traction loads on the pile shaft. In their publications no insight is given in the required magnitude of the additional loads, only a remark is made

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stressed before the actual pile bearing capacity calculation is performed [39]. The pre-stressing is performed by first loading the soil around the pile with prescribed displacements and subsequently resetting the displacements while preserving the stress state. Also in this method the magnitude of the prescribed displacements is not known a priori, rendering this method less suitable for predictions.

The large relative displacements generated when modelling the pile penetration as a circular cavity expansion or downward moving rigid element defeats almost all updated Lagrangian algorithms, as severe mesh distortion leads to numerical instability. The mesh distortions are partly overcome when the distorted elements are remeshed. The frequent remapping of all the variables, i.e. [93], on the new mesh is a cause of additional errors for this method. For pile installation the main areas where mesh distortion occurs are the pile base and shaft. This problem can be solved by incorporating a precut hole in the center of the mesh under the pile base [83], [54]. The hole allows for a horizontal displacement of the nodes on the symmetry axis. The hole widens as the pile is pushed in the soil. Also, they used a contact algorithm which updates or ‘remeshes’ the contact elements, making large deformations on the contact possible.

In contrast to total Lagrange and updated Lagrange methods the Eulerian and Arbitrary Lagrangian-Eulerian (ALE) schemes allow for uncoupling of the mesh and material. In the Eulerian method the mesh is fixed whereas in the ALE method the mesh and material can move separately. Element shapes can be independently optimized from deformations. Therefore no mesh distortions will occur, e.g., [79]. The ALE method is successfully im-plemented for the simulation of the cone penetration test by adding a Drucker-Prager and Mohr-Coulomb constitutive model within the ALE-framework [28], [29]. Although, strictly speaking these calculations were Eulerian (as the mesh is fixed). The Drucker-Prager con-stitutive model was used for the majority of the calculations, as it proved to be difficult to obtain stable solutions for the Mohr-Coulomb model. More recently similar calculations were made with ABAQUS, that incorporates the ALE scheme, for cone penetration in sand [175] and in clay [169]. In sand the Drucker-Prager model was used, whereas for clay the CAM-CLAY critical state model is used. More recently the explicit scheme of ABAQUS was used for the study of cone penetration in a perfectly plastic material [190]. As, in the explicit scheme groundwater cannot be incorporated, this resulted in the study of solely un-drained (Von Mises constitutive model) or un-drained behaviour (Drucker Prager constitutive model). An adaptive mesh was used, i.e. the re-meshing is implemented by an ALE scheme. The particle-in-cell or material point method is another numerical method to deal with large deformations in numerical codes [199], [200]. This is a special variety of the ALE method. In this scheme first the mesh is deformed in the Lagrangian phase. This allows for the calculation of the strain and stress increments. Secondly, in the convective phase, a new position of the computational mesh is chosen, and the velocity field is mapped from the body particles to the mesh nodes.

A method similar to [83] to model cone penetration is presented in [3], but without a separate pile geometry. An initial hole on the symmetry-axis of the axi-symmetric mesh is widened by prescribing horizontal and vertical displacements (vertical displacements are taken 85% of horizontal displacements). The analysis was performed with a commer-cial finite difference scheme (FLAC). For the soil the Mohr-Coulomb constitutive model

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was used, the model parameters are carefully selected for each considered case (calibration chamber tests with known stress and density conditions).

The Eulerian and ALE schemes offer the best solution for solving large deformations problems, no re-meshing or artificial holes in the mesh are needed to model the penetration process. Numerical frameworks capable of large deformation are readily available, only lacking more advanced constitutive models or proper treatment of groundwater flow. If these models are extended with more elaborate constitutive models for soil and coupled to a storage equation for the modelling of groundwater flow the method potentially offers a proper framework for the simulation of jacked pile installation. Experimental study of the change in soil properties near a displacement pile should improve the selection of a proper constitutive model. The calculated pile and soil response should be validated against such experimental results.

2.3

Effects of Pile Installation on Bearing Capacity and

Stiffness

In practice a difference in bearing capacity and stiffness between driven piles and jacked displacement piles is found. This difference results from the pile installation. In this Section the pile response after pile installation is compared for jacked and driven piles in order to obtain more insight into the direct link between the installation method and the bearing capacity. The pile tests performed after installation assesses the relative soil disturbance originating from the installation. These tests do not offer quantitative analysis of the change in soil properties near the pile as those are not measured before, after or monitored during the installation.

The effect of the pile installation method on the difference in static bearing capacity of jacked and driven piles,Qdrivenvs. Qjacked, is investigated for compression and tension

piles. The bearing capacity is derived from pile load tests on installed piles. A relative difference between two installation methods is studied, no additional information on the underlying mechanisms can be extracted from these tests. The bearing capacity of tension piles is mainly governed by the shaft resistance, whereas compression piles are mobilizing shaft and base resistance. If the pile is not instrumented to measure base resistance and shaft resistance separately, only general information on total pile head load is reported. The stiffness of a pile foundation, derived from the load-deflection behaviour, can also give additional information on the effect of the different pile installation methods on the response of a pile foundation.

2.3.1

Bearing Capacity

Numerous authors [6], [32], [76], [115], [136], [196], [205], [207] have studied the ratio

Qdriven/Qjackedin model tests and field tests.

It is important to note that these conclusions are only valid for static bearing capa-city. Under dynamic loads jacked piles appear more sensitive to decay of the shaft resist-ance compared with driven piles, which already endured a dynamic load during installation [196]. The amplitude of the dynamic load seems to influence the resulting bearing capacity

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Figure 2.1: Lateral stress variations during static tension load tests after monotonic, jacked, and pseudo-dynamic installation of piles; sensors positioned at a with the pile diameterB normalized heighth above the pile base

[71]. For cyclicly loaded tapered piles they found that jacked piles retained their increased stiffness and loading capacity under load cycles with an amplitude < 25% of the static compression capacity or less than< 75% the static uplift capacity.

When performing a jacked installation in-situ the pile is jacked into the soil in several strokes of 0.5 – 1 m. It is nearly impossible to monotonically install the pile in a single stroke. In model pile tests the latter is not only possible, but almost always the installation method used. A difference in load capacity exist when a model pile is installed in a single or in multiple strokes [115], see Figure 2.1, where also the development of lateral stress is plotted. Clearly the lateral stress depends on the installation method and is the highest at the intermediate (h/B =3; three pile diameters located from the base) instrument level. This also can explain the difference between both jacked and driven bearing capacities reported by the same authors when compared to the monotonically jacked results of the other references.

The tests performed by the various authors mentioned above and the ratioQdriven/Qjacked

they found are summarized in Table 2.1. This load is obtained from pile load tests with a limited pile head settlement. The value does not reflect a steady penetration resistance. When model tests were performed additional information about the test setup is given, e.g., the usage of a calibration chamber or the acceleration level of the centrifuge test. To com-pare the test condition, a normalized installation depthL/D is given, with L the installation depth of the pile andD the pile diameter. A short summary of the initial soil conditions of the test is included where soil conditions were varied (i.e. loose, dense or a range of confining pressuresσc).

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Author (F)ield or (M)odel L/D Soil conditions Test Type Qdriven/Qjacked

Al-Mhaidib M 9-43 Loose and Dense Tension Loose: 0.6

& Edil [6] 1g drained sand Dense: 0.2

Bonita et al. [32] M 38 Loose and Dense Compression Loose: 0.6

calibration chamber saturated sand Medium: 0.4

Foray et al. [76] M 15-18 dry sand, Compression Base: 0.8

calibration chamber Dr= 80 % Shaft: 0.5

Lehane & M 13 drained C and T 0.95 - 1

White [115] 50g silica sand (pseudo driven)

Nauroy & M 15 saturated Tension σc= 50 kPa: 2.4

Le Tirant [136] calibration chamber silica sand σc= 200 kPa: 0.9

σc= 400 kPa: 0.77

White & M 13 drained Compression 0.95 - 1

Lehane [196] 50g silica sand (pseudo driven)

Yalcin & M 16 dry sand Tension 0.45 - 0.55

Meyerhof [205] 1g (driven/buried)

Yang et al. [207] F 100-200 decomposed granite Compression Shaft: 0.5 - 0.7

1g Base: 6

Table 2.1: Influence of pile installation method on the ratio of the static bearing capacity of driven and jacked pilesQdriven/Qjacked, of piles with embedded lengthL and diameter D

Three tests series have been executed in drained sand. In those cases the influence of the water remaining in the pores (e.g. capillary suction) is not elaborated [6], [115], [196]. This may play a role as they measured during installation, and in the case of pseudo-driving the absence of excess pore water pressures can not be guaranteed. For the low loading rate of 2.5 mm/min of [6] on the other hand this has a minor influence on the results.

TheL/D ratio of the piles used in all the model tests is small. Compared with a typical Dutch situation, whereL/D = 50, model piles are more rigid. Several model tests were performed at low stress conditions, and are not comparable to slender piles with high stress levels and stress gradients along the pile.

From the table can be concluded that the majority of the tests showQdriven/Qjacked<

1, i.e., the static bearing capacity of jacked piles is higher than that of driven piles. Only a very limited amount of tests is considered, also no distinction is made between compression and tension piles. The higher jacked pile capacity is most likely caused by the stress build up in the soil without as much unloading cylces as during pile driving, this is also found for the pile stiffness (see next Section).

2.3.2

Pile Stiffness

Similar to the bearing capacity the initial stiffness of the pile response is also influenced by the pile installation method. A difference in stiffness is found between continuous jacking and staged jacking [209]. During centrifuge model pile experiments in sand they found that a 20 minute pause during pile jacking (in practice necessary for the installation of long

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the continuous installation. The initial stiffness of the pile response becomes less stiff when the pile is jacked in multiple stages. In addition jacked piles have higher shaft stiffness than driven piles, whereas driven piles have a stiffer base response [207]. For single jacked open ended tubular piles larger stiffness is found than similar piles situated in a pile group [62].

2.4

Stress Change

As seen in the aforementioned section each installation method yield different static bearing capacity. The soil state changes if a displacement pile is forced into the soil. Unfortunately the combined measurement of stress change and density change are scarce. This section focusses on the experimental investigations into soil stress change induced by pile installa-tion. The magnitude of this stress change and the effects on the pile bearing capacity will improve the understanding of the soil behaviour near a displacement pile during installation. Ideally, the development of the stress in the soil should be recorded during the pile in-stallation phase. In the past the stress distribution has been qualitatively monitored by using the photoelastic analogon in which the soil is substituted by crushed glass or glass beads [58], [68] [116] and [189]. Only the research published in [7] [8] also quantified the mag-nitude of the principal stress and strains near a model pile. Unfortunately the resolution in time and space is rather limited. Clear stress development cannot be derived from the res-ults. One of the main findings is the occurrence of non-coaxiality (the principal stress and strain directions do not coincide). This non-coaxiality has already been theoretically invest-igated [60]. The potential of this method resulted in the design of an improved photoelastic measurement setup described in Chapters 3 & 4 for the investigation of stress and density change during the pile installation stage.

The photoelastic method is not applicable for natural soils, because these are not bi-refringent. Therefore, in natural soils the measurement of stress evolution near the pile is limited to a point evolution. The measurement can be performed by transducers embed-ded in the soil or transducers build into the pile wall. In many research projects piles have been instrumented to measure base capacity and local shaft resistances. The results of such projects are reported in§2.4.1. The usefulness of these measurements, often executed after pile installation are similar to the comparisons on bearing capacity, although with some-what more detail in the analysis of the stress distribution along the pile. Tests in which the horizontal contact stress is measured during the pile installation or the subsequent test are summarized in§2.4.2. An overview of the reported published papers on stress change is given in§2.4.3.

2.4.1

Local shaft friction

Instrumented piles are often only instrumented to measure base load and axial load on in-termediate levels. In most setups the local shaft friction is derived from vertical stress measurements on two levels on the pile [15]. In literature some discrepancy is found in the shape of the shaft friction distribution along the pile. With increasing depth and therefore increasing in-situ vertical and horizontal stresses and increasing stiffness of the soil a more

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shaft friction

parabolic

exponential

depth

Figure 2.2: Two typical shaft friction profiles; exponentially increaseing with depth (left) and parabolic shaped with its maximum value located above the pile base (right)

or less exponential shaft friction distribution is expected in a homogeneous soil [9], [46], [75], [118], [139], [196]. This is the left curve sketched in Fig. 2.2.

However, a parabolic distribution (right curve in Fig. 2.2) of the local shaft friction is measured in centrifuge tests at large acceleration levels (60g - 100g, prototype pile length of 53 m) [15]. The maximum shaft friction in these tests clearly occurs above the pile base. Others only found a parabolic shaft resistance distribution for their tests performed in carbonate sand [105]. This indicates that a parabolic distribution may occur when particle crushing is significant. In the first tests the high stresses may have led to particle crushing and in the second tests a combination of high stresses and weaker particles may be the cause. If residual stresses [37] are neglected in the analysis of the pile load test also a parabolic shaft friction distribution is found, e.g., [185]. These residual stresses in the pile are caused by an equilibrium situation of shaft resistance on the upper part of the pile that prevents the pile from moving upwards and shaft resistance below the 0-plane and the base resistance facing upward trying to push the pile upward (elastic rebound of the soil). As a result the pile is not fully unloaded. Neglecting these stresses in the interpretation leads to erroneous interpretation of the instrument readings during a static pile load test [75].

Reliable measurement of the shaft friction is not trivial. During pile installation the instrumentation can be damaged or the sensitivity of the strain-gauge bridge can change. However, after installation it is next to impossible to repair or re-calibrate the instrument levels. It is not uncommon in field tests that all instrument levels are calibrated using the calibration data of the sensor located in the pile head. When dealing with model piles all in-strument levels can be separately calibrated before the pile installation and after the pile has been retrieved. Therefore these measurements are more reliable than field measurements.

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(a) Influence of embedment length on station-ary radial effective stress

(b) τ′

rz vsσ′r for compression and tension

tests at full depth

Figure 2.3: Field tests on jacked displacement piles [113]; sensors positioned at a with pile radiusR normalized height h above the pile base

2.4.2

Horizontal Contact Stresses

If the horizontal stress has been measured at all, it has been measured in discrete points along the pile (e.g. [106]).

Field tests performed on jacked displacement piles (D = 102 mm, L = 6 m) in dense quartz sand (with an interval between the strokes of 300 s) [113] presented the local shaft friction and also the horizontal (radial) stress acting on the pile during pile jacking. Only the stationary stressesσrsin the interval between jacking were reported. During jacking

the horizontal stress was lower. The effect of locked in stresses due to partial unloading in these jacking intervals is not elaborated. The stationary horizontal effective stresses during installation are given in Figure 2.3(a) for three instrument levels located at respectively 8, 28 and 50 pile radii from the pile base. On the vertical axis in Figure 2.3(a) the embedded depth of the instrument level is given, i.e., the stress measurements are plotted for fixed depths. Clearly a reduction of radial stress and local shaft friction as function of pile dis-placement is observed from the measurements. The stress paths from the pile load tests at intermediate depth and at full depth are given in Figure 2.3(b). From this Figure it becomes apparent that the peak value for the horizontal stress is larger then the stationary value. This is also the case for the average shear stress (derived from the total load and the pile base load). This difference could be due to stress rotation or interface slip dilation. However, it is unclear why this effects could not occur during pile installation. In a subsequent publica-tion empirical estimapublica-tions have been derived for the long-term equilibrium stress state after installation and the increase in radial effective stress during loading of the pile [112].

In addition to single piles the effect of interaction between two steel displacement piles (D = 102 mm, L = 6 m) is investigated [46]. During the installation of the second pile (at 4.5D center to center to the last pile) a large increase in the effective radial stress around the first pile is found, as shown in Figure 2.4(a). The increase is largest when the base of the second pile is at the instrument level of the first pile. After the base had passed, a gradual

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(a) Measurements of radial stresses and shear stresses on the first pile against location of the second pile(during installation of the second pile)

(b) Pile tests on the first pile (pile A) before and after installa-tion of pile B and the second pile (pile B) after installainstalla-tion

Figure 2.4: Influence of the installation of a second pile on an already installed pile [46]

decrease in effective radial stress is found. The end value of the effective radial stress is twice the initial value after installation of the first pile. The negative shear stresses during the approach of the pile base indicate a downward movement of the soil around the first pile. When the pile base of the second pile has passed the instrument level of the first pile the shear stresses become positive, indicating an upward movement of the soil. The instrument that did not face the second pile directly registered only a slight increase in negative shear stress after the pile base of the second pile is passed, indicating a different stress regime at the other side of the pile

After installation of the second pile the first pile gained 51% in shaft capacity and lost 43% in base capacity. The overall capacity increased by 19%; the reduced base capacity indicates an uplift of the pile. The second pile yielded the same total capacity as the first pile after the installation of the second pile. The initial stiffness response of the retested first pile increased, while at the same time the base stiffness decreased (see the pile tests in Figure 2.4(b)).

From these tests can be concluded that in pile groups the load distribution along the pile will be different from single piles. The pile will react more stiff if the shaft resistance is relatively large compared with the base resistance. The mobilization of the shaft resistance requires less pile-soil displacement than the mobilization of the base resistance. In inner piles in a pile group will most likely have less asymmetry in the shear load than found in these tests, as these piles have eight neighbouring piles.

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(a) Normalized horizontal stress during monotonic in-stallation (mean of four tests, error bars 1 std dev.)

(b) Reduction in horizontal stress between instruments during monotonic installa-tion ration of h/B = 1 – 9 (mean of four tests, error bars 1 std dev.)

(c) Stationary horizontal stress degradation at h/B = 1 during one way compres-sion tests

(d) Stationary horizontal stress degradation at h/B = 1 during two way cyclic compression load tests(cycles from installa-tion are incorporated)

Figure 2.5: The effects of different types of pile installation on horizontal stress evolution [196]; sensors positioned at a with the pile diameterB normalized height h above the pile base

The horizontal stresses acting on the pile during installation of a model pile (design in [106]) in sand in a geotechnical centrifuge are presented in [105]. The tests were done at several acceleration levels in a quartzitic sand (Leighton Buzzard sand) and show a linear increasing horizontal stress and local shaft friction with depth both measured during install-ation. The tests performed in a weaker carbonate sand (Dog’s Bay sand), however, show a parabolic development of the horizontal stress and shaft friction. The largest value was found above the pile base level.

These tests effectively have only one loading cycle, the pile is continuously jacked. The amount of loading cycles does influence the horizontal stresses on the pile [196]. The meas-ured normalized horizontal stress during monotonic installation is plotted in Figure2.5(a). The results indicate that the normalized horizontal stress remains more or less constant. Even when different instrument levels are compared only a slight reduction is found. This is visible in Figure 2.5(b) where the instrument levels at 3, 6 and 9 timesD from the pile base are plotted. No friction reduction in normalized horizontal stress is found. The stress normalization withqcfrom the penetration test, that includes installation effects, results in

the loss of installation effects in these presented results.

The influence of the installation method on the horizontal stress during and after pile installation is studied by comparing the horizontal stress response after a set of loading cycles. The results are shown in Figure 2.5(c), 2.5(d). The behaviour of the horizontal stress is a function of the amount of loading cycles, both during installation and pile load test. During one way loading again a constant value is reached after 30 cycles (Figure 2.5(c)). During two way cyclic loading the horizontal stress reduces to zero.

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condi-Figure 2.6: Lateral stress profiles before and at ultimate capacity in static load tests

tions in the lower 50 mm of the 180 mm high sample. The influence of the remaining pore water on the results is not elaborated. Therefore the effects of unsaturated soil behaviour cannot be excluded.

These results show that in these tests the jacked piles are more sensitive for subsequent cyclic loading, whereas driven piles have already reached their limiting value for the hori-zontal stress. This also explain the higher stiffness response of jacked as most likely more shaft resistance is mobilized due to the larger horizontal stresses near the pile.

The horizontal stress distribution during pile load tests are shown in Fig. 2.6 [115]. The maximum value for the horizontal stress is found above the pile base level. Also the jacked piles yield a larger increase in lateral stress during the compressive pile load test than monotonic and pseudo-dynamic installed piles. However, in tension the pseudo-dynamic tests performs similar to the jacked installation.

2.4.3

Overview

An overview of the influence of pile installation on the measured stress change in the pile, on the soil-pile contact during or after installation is presented in Table 2.2. Similar to Table 2.1, the test conditions are listed. Additionally the type of pile installation, the type of meas-urement, ad whether measurements were performed during installation or during the load test after installation, are given. When residual stresses are considered in the analysis of the load test this is shown in columnσres(not applicable for measurements during

install-ation). The main results deal with the shape of the shaft friction distribution (exponential or parabolic), and the change of local horizontal stress at a certain depth with an increase of pile displacement (column∆σh). The long term effects [13], [97], [117] [118] are also

(33)

res h

Altaee [9] F Driving Load Test Yes 39 n.a. Exponential

Axelsson [13] F Driving Load Test ?? 77 incr. Parabolic2

Balachowski [15] M1 Jacking Installation n.a. 10-33 n.a. Parabolic

Chow [46] F Jacking Inst. & LT n.a. 60 decr. Exponential

De Nicola [139] M1

Both Inst. & LT Yes ca. 10 decr. Exp.

Fellenius & Altaee [75] F & M Driving Load Test Yes ca. 30 n.a. Exponential

Jardine et al. [97] F Driving Load Test n.a. 20-40 incr.5

n.a

Klotz & Coop [105] M1

Jacking Installation n.a. 23 decr. Exp./Par.6

Lehane et al. [113] F Jacking Installation n.a. 60 decr. Exponential

Lehane & White [115] M1

Jacking Inst & LT n.a. 13 decr.4

Parabolic3

Leung et al. [117] [118] M1

Jacking Installation n.a. 12 decr. Exponential

Vesic [185] F Driving Load Test No 34 n.a. Parabolic

White [193] M Jacking Installation n.a. 10 decr.(τ ) n.a.

White & Lehane [196] M1

Both Installation n.a. 21 No Exponential

Table 2.2: Influence of pile installation on stress; 1model test in geo-centrifuge; 2from statnam test lateral earth pressure is exponential; 3distribution is given forσh instead of

shaft friction;4stationary measurements show decreasingσh, moving pile shows increasing

σh;5increase in bearing capacity due to pile setup will be elaborated in Section 2.7; ;6silica

is exponential and carbonated is parabolical

2.5

Density Change

Only a limited amount of publications from literature which deal with the direct measure-ment of the volume change or density change due to pile installation are identified. For the measurements of the soil density three types of methods are used: optical methods (stereo-photography, PIV, X-ray), measurements of thermal or electrical resistance and lastly nuc-lear methods. The first method generally is more accurate, only except X-ray, an unobstruc-ted view is necessary. Therefore, the model reduces to a plane strain setup. Although the thermal or electrical probes are less accurate, an undisturbed true three dimensional soil fail-ure zone can develop around the pile. Nuclear back scattering methods are primarily used for soil characterization and not for the investigation of pile installation. The results from literature are separated into plane strain observations and three dimensional observations, starting with the latter.

2.5.1

Three Dimensional

With a X-ray method a close look at the soil behaviour around a jacked pile can be given [152]. The tests are done at low stress conditions in dry sand. The displacement of the soil is measured with radiography equipment: Cobalt-60 radioactive source and lead shot markers. Apart from the lead shot no additional disturbance is induced by the measurement method, the soil is free to flow on all sides of the pile. The soil below the pile base

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