Modelling Pile Installation Effects: A Numerical Approach

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(1)196807-os-Engin_DEF 1.indd 1. 23-5-13 17:43.

(2) PROPOSITIONS accompanying the thesis Modelling Installation Effects - A Numerical Approach by Harun K¨ ur¸sat Engin, Delft, 10 June 2013 1. Load settlement behaviour of displacement piles cannot be modelled properly without considering the installation effects. 2. The Press-Replace technique is a simple and stable method that can be used to model pile penetration problems in standard finite element software. 3. The strength of interface elements ought to be directly correlated to the strength of continuum elements. 4. The stresses around a wished-in-place pile that are imposed to approximate the installation effects, reduce during the nil step in which the unbalance is removed. This reduction can be minimised by using a stiff material when imposing the stresses and restoring equilibrium. 5. The accuracy at which BVPs such as pile penetration (which involve complicated stress paths) can be simulated, is not guaranteed by tuning the constitutive model only for well-defined stress/strain paths. 6. The relative density, Id is not a reliable input parameter; however, if it could be determined accurately, it is well suited for estimating the properties of cohesionless soils. 7. The pleasure of doing research stands on the motivation to overcome frustration and to obtain the final result. 8. When a novelty is introduced, emphasizing the capabilities should not supress the discussion on the limitations. 9. Engineers and managers are alike; the former use methods, the latter use people that, in both cases, have not been developed by themselves. 10. The current scientific publishing system shows how academia is exploited twice: by letting articles be reviewed for free and successively asking money to access the published articles. Since the review process is done without any cost, there should be no concern for the quality of open-access articles. 11. An overdose of agitation makes people indifferent. These propositions are regarded as opposable and defendable, and have been approved as such by the supervisors Prof.ir.A.F. van Tol and Dr.ir.R.B.J. Brinkgreve.. 196807-st-Engin.indd 1. 22-5-13 16:17.

(3) STELLINGEN behorende bij het proefschrift Modelling Installation Effects - A Numerical Approach van Harun K¨ ur¸sat Engin, Delft, 10 Juni 2013 1. Last-verplaatsingsgedrag van grondverdringende palen kan niet goed worden gemodelleerd zonder de installatie effecten te beschouwen. 2. De Press-Replace techniek is een eenvoudige en stabiele methode die kan worden gebruikt om het inbrengen van palen in standaard eindige-elementen software te modelleren. 3. De sterkte van interface elementen dient direct gecorreleerd te worden aan de sterkte van continu¨ um elementen. 4. De spanningen rondom een wished-in-place paal die worden aangebracht om de installatie effecten te benaderen, nemen af tijdens de nil step waarin de onbalans wordt verwijderd. Deze afname kan worden geminimaliseerd door gebruik te maken van stijf materiaal tijdens het aanbrengen van de spanningen en het herstellen van evenwicht. 5. De nauwkeurigheid waarmee randvoorwaardeproblemen zoals het inbrengen van palen (hetgeen gecompliceerde spanningspaden met zich mee brengt) kan worden gesimuleerd, is niet gegarandeerd door het constitutieve model alleen af te stemmen op goed-gedefinieerde spanning/rek paden. 6. De relatieve dichtheid, Id , is geen betrouwbare invoer parameter; echter, als deze nauwkeurig zou kunnen worden bepaald, is deze zeer geschikt om de eigenschappen van cohesieloze gronden te schatten. 7. Het plezier aan het doen van onderzoek staat met de motivatie om frustratie te overwinnen en het eindresultaat te behalen. 8. Wanneer een noviteit wordt ge¨ıntroduceerd dient het benadrukken van de mogelijkheden de discussie over de beperkingen niet te onderdrukken. 9. Ingenieurs en managers zijn vergelijkbaar; de eersten gebruiken methoden, de laatsten gebruiken mensen die, in beide gevallen, niet door hen zelf zijn ontwikkeld (gevormd). 10. Het huidige wetenschappelijk publicatie system laat zien hoe de academische wereld dubbel wordt uitgebuit: door onbetaald artikelen te laten reviewen en door vervolgens te laten betalen om toegang te verkrijgen tot de gepubliceerde artikelen. Omdat het reviewen toch onbetaald is, zijn zorgen over de kwaliteit van artikelen in vrij toegankelijke journals, ongegrond. 11. Een overdosis aan onrust maakt mensen onverschillig. Deze stellingen worden opponeerbaar en verdedigbaar geacht en zijn als zodanig goedgekeurd door de promotoren Prof.ir.A.F. van Tol en Dr.ir.R.B.J. Brinkgreve.. 196807-st-Engin.indd 2. 22-5-13 16:17.

(4) Modelling Pile Installation Effects A Numerical Approach.

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(6) Modelling Pile Installation Effects A Numerical Approach. PROEFSCHRIFT ter verkrijging van de graad van doctor aan de Technische Universiteit Delft, op gezag van de Rector Magnificus prof.ir. K.C.A.M. Luyben, voorzitter van het College voor Promoties, in het openbaar te verdedigen op maandag 10 juni 2013 om 12.30 uur ˙ door Harun K¨ ur¸sat ENGIN Master of Science in Civil Engineering Middle East Technical University, Ankara, Turkey geboren te Ankara (Turkije)..

(7) Dit proefschrift is goedgekeurd door de promotoren: Prof. ir. A.F. van Tol Copromotor: Dr. ir. R.B.J. Brinkgreve. Samenstelling promotiecommissie: Rector Magnificus, voorzitter Prof. ir. A.F. van Tol, Technische Universiteit Delft, promotor Dr. ir. R.B.J. Brinkgreve, Technische Universiteit Delft, copromotor Prof. Dr. M.A. Hicks, Technische Universiteit Delft Prof. Dr. ir. L.J. Sluys, Technische Universiteit Delft Prof. Dr. ing. habil. I. Herle, Technische Universität Dresden Prof. Dr. ing. H.P. Jostad, Norges Teknisk-Naturvitenskapelige Universitet Dr. L. Zdravković, Imperial College London. This research is supported by the Dutch Technology Foundation STW, which is part of the Netherlands Organisation for Scientific Research (NWO) and partly funded by the Ministry of Economic Affairs (project number 10189). The research leading to these results has also received funding from the 7th Framework Programme (FP7/2007-2013 under grant agreement PIAG-GA-2009-230638). The findings reflect only the authors views and the EC is not liable for any use that may be made of the information contained therein.. ISBN 978-94-6186-140-5. © 2013 by H.K. Engin. All rights reserved. No part of the material protected by. this copyright notice may be reproduced or utilized in any form or by any means, electronic or mechanical, including photocopying, recording or by any information storage and retrieval system, without written consent of the publisher.. Cover: Erjona Engin & H. K¨ ur¸sat Engin (Concept); A. Burak Veyiso˘glu (Design) Printed by : Ipskamp Drukkers / E-mail: info@ipskampdrukkers.nl.

(8) ˙ To my grandfather Ibrahim Hakkı Bilici (RIP) & all others who have supported me.

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(10) Summary One of the most traditional methods for supporting structures resting on soft soils is the use of piles. They generally work by transferring the loads to deeper soil layers, which can provide sufficient bearing capacity when mobilised. This type of foundations has been commonly used throughout the world and also in the western part of the Netherlands due to the typical soil profile which consists of a thick (10 20m) soft soil layer underlain with a stiff bearing stratum composed mostly of quartz sand. In this perspective, this study sheds a light on the behaviour of piles installed in silica sand. Piles can be installed in different ways. The most common way of installing the prefabricated piles is by using an impact hammer. With this driving method, the pile is penetrated by each drop or blow of the ram until the desired penetration level is achieved. Another alternative is the vibratory driving technique where the pile is forced to penetrate by a heavy vibratory head on top. Recent developments have allowed the piles to be jacked thus resulting in reduced nuisance and in higher capacity and stiffness when compared to the hammering and vibratory driving techniques. Yet, the method is generally limited to open ended piles or piles with relatively small diameters for closed ended piles. The bearing capacity of a pile depends on the soil properties and the stress state it is surrounded with. This is because the behaviour of granular material is governed by the packing of the grains and the contact stresses in between. The mean stress and the density can be described as the soil state, and the soil behaviour is determined on the basis of this state and the loading conditions. In the case of a displacement pile, the installation process causes a considerable amount of soil displacement and high levels of (reaction) stresses. These effects of pile installation are transmitted to soil through the interaction between sand grains and the pile, resulting in an altered soil state and properties. It is therefore obvious that one should have different installation effects for different installation techniques. For example, the confining stresses resulting from the dynamic installation technique (hammering or vibratory driving) are lower than those of the jacking technique. Depending on the density, dilation would be expected near the pile shaft during jacking. On the other hand, dynamic installation methods cause compaction near the pile. The installation effects may have two important consequences in the form of the influence on the performance of the pile (e.g. load displacement behaviour, capacity) in its service life and the influence (e.g. noise, vibrations) on the neighbourhood /environment. vii.

(11) viii. Summary. A more realistic behaviour and therefore an improved design would be achieved by considering the installation effects in the analyses than performing the analysis considering a geostatic stress state around the pile(s) modelled. In current practice, the installation effects are taken into account by some empirical design methods in order to estimate the bearing capacity of foundation piles. Several field and model tests performed to investigate the influence of pile installation on the bearing capacity, have led to an evolution of the empirical models to estimate the bearing capacity of displacement piles. Recent attempts to investigate the change in the soil state were also limited either to the measurement domain (generally close to the pile) or resolution as well as the variables (e.g. displacement, strain, stress, density) that can be quantified. However, the behaviour of piles during installation, the interaction with the surrounding soil and the resulting alteration of soil properties during installation are still not well known. This information is essential, not only to make better predictions of the pile bearing capacity and its behaviour in the soil under different loading conditions, but also to be able to predict the (side) effects of pile installation on the neighbourhood. The increase in the computational power allowed improved numerical techniques like the finite element method (FEM); however the installation effects of displacement piles have not been incorporated in the numerical models for engineering purposes. Because the installation effects have been discarded there will be considerable difference between the FE predictions and the real behaviour (performance) of the pile. Therefore it is important to model the installation effects of displacement piles numerically so that the load displacement behaviour of the pile during its service life as well as the secondary effects on neighbouring structures can be predicted more appropriately. For an accurate analysis of the pile installation and its consequent effects, all important aspects (e.g. large deformation effects, pore pressure effects, inertia effects and use of a suitable constitutive model) should be included in the numerical model. Nevertheless, simplifications are unavoidable. For example the inertia effects can be discarded and drained conditions (no excess pore pressure effects) can be assumed to hold for the pile jacking case, which is seen as quasi-static penetration. Therefore in the resulting FE model, the level of complexity will be reduced. This study is the numerical part of a larger project in which the installation effects of driven piles are investigated, both experimentally and numerically. The objective of this numerical study is to investigate and model the installation effects of pile jacking in sand in a numerical framework. Since any jacking operation can be considered as a quasi- static loading operation, the dynamic effects are not included in the analyses. For the same reason, drained conditions are assumed. As a common tool used in engineering practice, FE modelling is selected as the numerical method. In order to model pile jacking using the FEM in a standard FE code, some additional aspects have to be considered to account for the large deformation effects (e.g. change of geometry due to pile penetration). Bearing in mind the aforementioned points and the ease of applicability in a standard FE package, the Press-Replace (PR) technique was introduced. The technique benefits the use of a simple and robust small deformation formulation on one side and models the large deforma-.

(12) ix tion effects (e.g. the penetration process) by using updated geometry on the other side. The capabilities and limitations of the method were investigated by means of a parametric study. Based on this study, an optimised modelling technique was proposed. Another important aspect of FE analyses is that of modelling the soil behaviour. First of all, the stress and density dependency is an important feature of the soils to be modelled. This feature is quite well established by the Critical State theory in which the state (stress and density) determines the soil response. There are several Critical State models mentioned in the literature. Hypoplasticity was employed as the constitutive model in the simulation considering its capabilities (e.g. the void ratio and stress dependent stiffness). It has been previously observed that due to the high stress levels encountered at the pile base, particle or grain crushing may also occur. As a result of crushing, not only the behaviour but more importantly the material properties change. The constitutive model should therefore account for this modification. In order to model the crushing effects, the hypoplastic model was modified. In the modification the descriptive input parameters were allowed to alter during the simulations. However, this gave rise to problems especially in regions with high gradients of stresses. As a result, a converged solution with the modified version could not be obtained. In order to keep the model simple and still applicable in a standard FE code, the grain crushing effect was ignored in the FE simulations. As the aim of the study was to model installation effects of piles, the PR technique was employed to investigate the installation effects for different pile geometries and different soil densities. The results have shown that the installation field is qualitatively similar for all of these variations. This makes the generalisation of all these installation fields possible. In view of the results obtained from FE simulations, the possibilities of incorporating the installation fields (e.g. Cartesian stress and void ratio distributions) around a jacked pile were investigated. The installation fields were mapped to different meshes to investigate the possibility of imposing the installation fields. The idea was to account for the installation effects without simulating the penetration process and, as a result, model the service (e.g. load -displacement) behaviour of the displacement pile properly. The results have shown that the axial load displacement response of the pile at the end of the PR analysis was successfully captured by the imposed installation field around a wished-in-place pile modelled with a different discretisation than the PR model. For ease of applicability in engineering practice, the possibility of describing the installation effects in a mathematical (functional) form was investigated. This allows a versatile method for imposing the installation field around a wished-in-place pile so that it captures the jacked pile behaviour. Firstly a multiple regression algorithm was employed to fit functions for the Cartesian stresses and the void ratio for a reference case. The exponential form of the functions closely fit the installation surfaces. Secondly, these functions were extended to include the pile geometry and relative density effects..

(13) x. Summary. Despite the limitations and simplifications, it was shown that the idea of describing the installation field in terms of functional forms works reasonably well and can easily be applied in a standard FE analysis. In order to validate the functions proposed, a centrifuge pile was modelled. As a result of incorporating the installation effects around the wished-in-place pile, the load displacement response could be modelled reasonably well. In practice, if the installation effects are considered in the FE analyses, better insight into foundation design can be obtained. As a result of considering increased capacity and stiffness of the displacement pile analysed, the foundation costs would be reduced. Furthermore, if there are any buildings in the vicinity of a pile, possible disturbance or damage to that structure could be foreseen in the FE analysis..

(14) Samenvatting (in Dutch) Een van de meest traditionele methoden voor het ondersteunen van constructies op slappe grond is het gebruik van palen. Die werken in het algemeen door de belasting naar dieper gelegen grondlagen over te brengen, waardoor voldoende draagkracht kan worden gemobiliseerd. Dit soort funderingen is veelvuldig toegepast over de hele wereld en zo ook in het westen van Nederland door het typische grondprofiel met een dikke (10 - 20 m) slappe grondlaag met daaronder een stijve draagkrachtige laag bestaande uit voornamelijk kwarts zand. Vanuit dat perspectief bezien is deze studie bedoeld om inzicht te geven over het gedrag van palen in silicium zanden. Palen kunnen op verschillende manieren worden ge¨ınstalleerd. De meest gebruikelijke manier om prefab palen te installeren is door middel van een heiblok. Met deze inbrengmethode wordt de paal bij elke slag een stukje ingebracht totdat het vereiste inheiniveau is bereikt. Een alternatief is de trilmethode waarbij de paal de grond wordt ingebracht door een zwaar trilblok bovenop de paal. Recente ontwikkelingen hebben geleid tot het in de grond drukken van de paal waarbij overlast worden beperkt en draagkracht en stijfheid hoger (kunnen) zijn vergeleken met de heien trilmethode. Toch wordt de methode nog beperkt tot open buispalen of massieve palen met een relatief kleine diameter. De draagkracht van een paal hangt af van de grondeigenschappen en de spanningstoestand rondom de paal. Dat komt omdat het gedrag van granulair materiaal wordt bepaald door de pakking van de korrels en de contactspanningen ertussen. De gemiddelde spanning en de dichtheid kunnen worden aangegeven als de toestand van de grond, en het grondgedrag wordt bepaald op basis van deze toestand en de belastingsituatie. Bij een grondverdringende paal veroorzaakt het installatieproces een aanzienlijke grondverplaatsing en hoge (reactie-)spanningen. Deze paal installatie effecten worden overgebracht naar de grond door middel van de interactie tussen de zandkorrels en de paal, waardoor de toestand van de grond en eigenschappen wijzigen. Het is duidelijk dat verschillende installatietechnieken tot verschillende installatie effecten leiden. Bijvoorbeeld, de spanningen rondom de paal die voortkomen uit dynamische installatiemethoden (heien of trillen) zijn kleiner dan bij het in de grond drukken. Afhankelijk van de dichtheid wordt dilatantie verwacht langs de paalschacht tijdens het indrukken. Daar staat tegenover dat dynamische installatiemethoden compactie langs de paal veroorzaken. De installatie effecten kunnen twee belangrijke consequenties hebben in de zin van invloed op het gedrag van de paal (t.w. last-zakkingsgedrag, draagkracht) in de gebruiksfase en de invloed op de xi.

(15) xii. Samenvatting (in Dutch). omgeving (o.a. lawaai, trillingen). Een meer realistisch gedrag en derhalve een verbeterd ontwerp kan worden bereikt door de installatie effecten in de analyse mee in beschouwing te nemen ten opzichte van het modelleren van een zuiver geostatische spanningstoestand rondom de paal (palen). In de huidige praktijk worden installatie effecten meegenomen in empirische ontwerpmethoden ter bepaling van de draagkracht van funderingspalen. Diverse veld- en model testen, uitgevoerd om de invloed van de installatie op de draagkracht van de paal te onderzoeken, hebben geleid tot veranderingen in empirische modellen ter bepaling van de draagkracht van grondverdringende palen. Recente pogingen om de verandering van de toestand van de grond te onderzoeken waren beperkt in het meetgebied (in het algemeen rondom de paal) of de meetpuntdichtheid, als ook het aantal gemeten variabelen (verplaatsing, rek, spanning, dichtheid). Echter, het gedrag van palen tijdens installatie, de interactie met de omliggende grond en de resulterende verandering van de grondeigenschappen tijdens installatie zijn nog steeds niet goed bekend. Deze informatie is essentieel, niet alleen om betere voorspellingen te maken van de paal draagkracht en het gedrag in de grond onder verschillende belastingcondities, maar ook om de (neven-)effecten van de paal installatie op de omgeving te kunnen voorspellen. De toename in rekenkracht van computers heeft het gebruik van geavanceerde numerieke methoden zoals de eindige-elementenmethode (EEM) mogelijk gemaakt; echter de installatie effecten van grondverdringende palen zijn nog niet meegenomen in numerieke modellen voor ontwerpdoeleinden. Vanwege het negeren van installatie effecten is er een aanzienlijk verschil tussen de EEM predicties en het werkelijke gedrag van de paal. Daarom is het belangrijk om de installatie effecten van grondverdringende palen numeriek te modelleren zodat het last-verplaatsingsgedrag van de paal tijdens de gebruiksfase zowel als de secundaire effecten op omliggende constructies beter kunnen worden voorspeld. Voor een nauwkeurige analyse van de paalinstallatie en daaruit voortkomende effecten is het van belang om alle belangrijke aspecten (zoals grote deformatie effecten, waterspanningen, massa traagheid en het gebruik van een geschikt constitutief model) mee te nemen in het numerieke model. Desalniettemin zijn vereenvoudigingen onvermijdelijk. Daarom wordt uitgegaan van in de grond gedrukte palen. Traagheidseffecten kunnen dan bijvoorbeeld worden weggelaten en gedraineerde condities (geen wateroverspanningseffecten) kunnen worden verondersteld te gelden, hetgeen wordt gezien als quasi-statische inbrenging. Daardoor zal de complexiteit in het resulterende EEM model worden beperkt. Deze studie is het numerieke deel van een groter project waarin installatie effecten van geheide palen worden onderzocht, zowel experimenteel als numeriek. Het doel van deze numerieke studie is om de installatie effecten van in de grond gedrukte palen in zand in een numerieke context te onderzoeken en te modelleren. Omdat elke indrukking als een quasi-statische belasting kan worden beschouwd worden dynamische effecten niet meegenomen in de berekeningen. Om dezelfde reden worden gedraineerde condities aangenomen. Als een veelgebruikt gereedschap in de ingenieurspraktijk wordt de EEM modellering als numerieke methode gehanteerd..

(16) xiii Om het indrukken van palen met de EEM in een standaard programma te modelleren moeten enkele aanvullende aspecten worden beschouwd om de grote deformatie effecten in rekening te brengen (t.w. de verandering van de geometrie als gevolg van het inbrengen van de paal). Vanuit deze gedachte en de eenvoudige toepasbaarheid in een standaard EEM pakket is de Press-Replace (PR) techniek geintroduceerd. Deze techniek laat enerzijds het gebruik van een eenvoudige en robuuste kleine-deformatie formulering toe en modelleert anderzijds de grote vervormingseffecten (het inbrengproces) door gebruik te maken van geometrie- aanpassing. De mogelijkheden en beperkingen van de methode zijn onderzocht door middel van een parameterstudie. Op basis van deze studie is een geoptimaliseerde modelleringstechniek voorgesteld. Een ander belangrijk aspect van de EEM analyses is de modellering van het grondgedrag. Ten eerste is de spannings- en dichtheidsafhankelijkheid een belangrijke eigenschap van de te modelleren gronden. Deze eigenschap is redelijk verdisconteerd in de Critical State theorie waarbij de toestand (spanning en dichtheid) de grondrespons bepaalt. Er worden diverse Critical State modellen genoemd in de literatuur. Hypoplasticiteit is hier gebruik als constitutief model in de simulaties vanwege zijn eigenschappen (t.w. de spanningen poriëngetal-afhankelijkheid van de stijfheid). Het is eerder waargenomen dat ten gevolge van hoge spanningsnivieaus aan de paalvoet deeltjes of korrels kunnen verbrijzelen. Als gevolg van verbrijzeling verandert niet allen het gedrag maar veranderen ook de materiaaleigenschappen. Het constitutieve model moet daarom rekening houden met deze verandering. Om de verbrijzelingseffecten te modelleren is het hypoplastische model aangepast. In het aangepaste model mogen de modelparameters tijdens de simulatie veranderen. Echter dit gaf aanleiding tot problemen, in het bijzonder in de gebieden met hoge spanningsgradiënten. Als gevolg daarvan kon een geconvergeerde oplossing met het aangepaste model niet worden verkregen. Om het model eenvoudig en toepasbaar te houden in een standaard EEM programma is het verbrijzelen van korrels in de EEM simulaties verder buiten beschouwing gelaten. Met oog op het doel van deze studie om installatie effecten van palen te modelleren is de PR techniek toegepast om de installatie effecten bij verschillende paal geometrieën en verschillende dichtheden te onderzoeken. De resultaten hebben laten zien dat de verdeling van installatie effecten kwalitatief gelijk is voor al deze variaties. Dit maakt het mogelijk om deze effecten te generaliseren. In het licht van de met de EEM simulaties verkregen resultaten zijn de mogelijkheden om de installatie effecten (t.w. Cartesische spanningen en poriëngetal verdeling) rondom een in de grond gedrukte paal mee te nemen, onderzocht. De verdeling van installatie effecten is op verschillende meshes geprojecteerd om de mogelijkheden voor het aanbrengen van deze effecten te onderzoeken. Het idee is om de installatie effecten in rekening te brengen zonder het inbrengproces te simuleren, met als resultaat om het (last-verplaatsings-)gedrag van een grondverdringende paal in de gebruiksfase juist weer te geven. De resultaten laten zien dat de axiale lastverplaatsingsrespons aan het eind van de PR analyse succesvol wordt weergegeven met de aangebrachte installatie effecten rondom een zogenaamde wished-in-place paal gemodelleerd met een andere discretisatie dan het PR model..

(17) xiv. Samenvatting (in Dutch). Voor een eenvoudige toepassing in de ingenieurspraktijk is de mogelijkheid onderzocht om de installatie effecten in een wiskundige vorm (functie) weer te geven. Dit zorgt voor een algemeen toepasbare methode voor het aanbrengen van installatie effecten rondom een wished-in-place paal die het gedrag van een in de grond gedrukte paal juist weergeeft. In eerste instantie is een meervoudig regressie algoritme toegepast ter benadering van de functies voor de Cartesische spanningen en het poringetal voor een referentie situatie. De exponentiële vorm van de functies sluit nauw aan bij de verdelingen. Vervolgens zijn deze functies uitgebreid met de paal geometrie en de relatieve dichtheidseffecten. Ondanks de beperkingen en vereenvoudigingen is aangetoond dat het idee om installatie effecten te beschrijven in functie-vorm redelijk goed werkt en eenvoudig kan worden toegepast in standaard EEM berekeningen. Om de voorgestelde functies te valideren is een centrifugetest gemodelleerd. Als gevolg van het meenemen van de installatie effecten rondom een wished-in-place paal kon het last-verplaatsingsgedrag redelijk worden gemodelleerd. Wanneer in de praktijk de installatie effecten in EEM berekeningen worden meegenomen kan een beter inzicht in het funderingsontwerp worden verkregen. Als gevolg van de toegenomen draagkracht en stijfheid van de geanalyseerde grondverdringende paal kunnen de funderingskosten worden gereduceerd. Verder, als er gebouwen in de omgeving van de paal aanwezig zijn, kan mogelijke verstoring van of schade aan de constructie worden voorzien middels de EEM analyse..

(18) Contents Summary. vii. Samenvatting (in Dutch). xi. List of Symbols. xxix. 1 Introduction 1.1 Background . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.2 Scope and Limitations . . . . . . . . . . . . . . . . . . . . . . . . . . 1.3 Outline . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 1 1 4 4. 2 Background on Pile Installation 2.1 Introduction . . . . . . . . . . . . . . . . . . . . . . 2.2 Estimation of Bearing Capacity . . . . . . . . . . . 2.2.1 Empirical Correlations . . . . . . . . . . . . 2.2.2 Cavity Expansion and Strain Path Methods 2.2.3 Numerical Analysis . . . . . . . . . . . . . . 2.3 Investigation of Installation Effects . . . . . . . . . 2.3.1 Stress Change . . . . . . . . . . . . . . . . . 2.3.2 Density Change . . . . . . . . . . . . . . . . 2.3.3 Material Change . . . . . . . . . . . . . . . 2.3.4 Effects on Pile Behaviour . . . . . . . . . . 2.4 Including Installation Effects in the Analyses . . . 2.4.1 Empirical Methods . . . . . . . . . . . . . . 2.4.2 Numerical Methods . . . . . . . . . . . . . 2.5 Conclusions . . . . . . . . . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. 7 7 7 8 9 11 16 16 18 23 26 31 31 33 34. 3 Constitutive Modelling 3.1 Introduction . . . . . . . . . 3.2 Hypoplasticity . . . . . . . 3.3 Hypoplastic Grain Crushing 3.4 Conclusions . . . . . . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. 37 37 38 42 47. . . . . . . . . Model . . . .. xv. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . ..

(19) xvi 4 Press-Replace Technique 4.1 Introduction . . . . . . . . . . . . . . . . . . 4.2 Technique . . . . . . . . . . . . . . . . . . . 4.3 Validation . . . . . . . . . . . . . . . . . . . 4.3.1 Cone Penetration in Undrained Clay 4.3.2 Pile Penetration in Undrained Clay . 4.4 Parametric study . . . . . . . . . . . . . . . 4.4.1 Element Type and Mesh Density . . 4.4.2 Step Size . . . . . . . . . . . . . . . 4.4.3 Interface Properties . . . . . . . . . 4.5 Discussion and Conclusions . . . . . . . . .. Contents. . . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . .. 51 51 52 58 58 63 64 66 67 70 76. 5 Investigation of Installation Effects 5.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . 5.2 Numerical Modelling - Application of the PR Technique 5.3 Results of the FE Simulations . . . . . . . . . . . . . . . 5.4 Discussion of Results and Conclusions . . . . . . . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. 79 79 80 83 92. . . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . .. 6 Generalisation of Installation Effects 6.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2 Approximation of the Installation Effects by Interpolation . . . . . . 6.2.1 Method - Triangular Mesh Interpolation . . . . . . . . . . . . 6.2.2 Results and Discussions . . . . . . . . . . . . . . . . . . . . . 6.2.3 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3 Approximation of the Installation Effects by Bivariate Nonlinear Regression . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.1 Theoretical Background . . . . . . . . . . . . . . . . . . . . . 6.3.2 Describing Installation Effects by Surface Fitting . . . . . . . 6.3.3 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.4 Discussion of Results and Conclusions . . . . . . . . . . . . . 6.4 Generalisation of the Installation Effects . . . . . . . . . . . . . . . . 6.4.1 Determination of the Coefficients of Generalised Functions . . 6.4.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.5 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 109 109 110 115 125 125 126 126 130. 7 Application 7.1 Introduction . . . . . . . . . . . . . . . . . . 7.2 Verification Cases . . . . . . . . . . . . . . . 7.3 Validation Case - Centrifuge Pile Jacking in 7.4 Discussion of Results and Conclusions . . .. 133 133 133 136 143. . . . . . . . . . . . . . . . . Baskarp Sand . . . . . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. . . . .. 97 97 98 98 105 108. 8 Conclusions 145 8.1 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 145 8.2 Outlook for Further Research . . . . . . . . . . . . . . . . . . . . . . 150 Bibliography. 152.

(20) Contents. xvii. A Calculation of Unbalance. 167. B Implementation of Functions. 171. C Calculation of Pile Bearing Capacity Using Empirical Methods. 173. Acknowledgements. 185. Curriculum Vitae. 187. Publications. 189.

(21) xviii. Contents.

(22) List of Figures 2.1. a. Definition of normalising parameter pcs ; Normalisation of the b. base and c. shaft bearing capacity factors for Leighthon Buzzard sand (after [90]) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2 Comparison of MPM and ALE methods: 1. Beginning of calculation step in which the position, velocity, current strain and stress, state variables and local element coordinates are known at time t 2. Mapping of particle velocities to nodes 3. (Updated) Lagrangian phase in which the element is deformed and 4. the position, velocity and strain, stress and state variables are updated at the material points. 5.a. The Lagrangian mesh is reset. 5.b. The mesh is updated arbitrarily (after [27]) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3 Large deformation (UL) analyses using frictional contact, where smooth geometries for pile tip are used a. [63] b. [112] c. [77] . . . . . . . . . 2.4 Average horizontal and vertical stress distribution based on photoelastic measurements ([56]) . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.5 Radial stress distribution after 7 m of continuous pile jacking in a. dense and b. loose Karlsruhe sand; c. radial stress distribution at 4m depth (after [112]) . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.6 Radial stress distribution of pile hammering in a. dense and b. loose Karlsruhe sand; c. radial stress distribution at 3 m depth (after [112]) 2.7 Radial stress distribution of vibratory pile driving in a. dense and b. loose Karlsruhe sand; c. radial stress distribution at 3 m depth (after [112]) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.8 Change in porosity during pile penetration for different initial conditions at different measurement locations (after [56]) . . . . . . . . . . 2.9 a. Test setup for the centrifuge pile tests; b. The sketch of the density measurement principle, where the potential drop was measured between the conductors; the continuous lines are the electric field and the arrows show the current flow. . . . . . . . . . . . . . . . . . . . . 2.10 Displacement and volumetric observations in 1 g models using a. radiograph measurement ([128]), b. thermal conductivity ([38]) and c. imaging techniques (in crushable carbonate sand, [180]) . . . . . . . 2.11 Void ratio distribution after pile penetration simulations based on a. FE with remeshing [78] and b. CEL [124] in medium dense Karlsruhe and Mai-Liao sands, respectively . . . . . . . . . . . . . . . . . . . . xix. 9. 15 15 16. 17 17. 17 19. 20. 20. 21.

(23) xx. List of Figures 2.12 Void ratio distributions of medium-dense (Id = 0.45) Karlsruhe sand at several distances from the pile shaft for a. jacking, b. hammering and c. vibratory driving installation methods ([112]) . . . . . . . . . 2.13 Grain crushing for flat pile base a. [164] b. [103] and for c. conical base [103] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.14 Results of DEM simulation a. without and b. with particle breakage [104] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.15 DEM results incorporating particle crushing for different pile types [105] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.16 Idealisation of induced tensile stress in a particle and the crushing mechanism [105] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.17 A general critical state line for sands [132] . . . . . . . . . . . . . . . 2.18 a. End bearing mobilisation, and b. shaft friction transfer curves of displacement and non-displacement piles obtained at different segments of on the pile for a medium dense FF sand (after [45]) . . . . 2.19 Load transfer (t-z) curves for a. simplified elastic-perfectly plastic soil and b. for strain hardening and softening soil (after [125]); dashed lines represent the confinement and the material change effects. Mobilisation of shear strength at different h/R ratios for c. simplified elastic-perfectly plastic soil and d. for strain hardening and softening soil. Field observations on the mobilisation of e. shear strength mobilisation and f. shaft friction [98] . . . . . . . . . . . . . . . . . . 2.20 a. Typical shaft friction distributions [56]; b. Residual loads during and after installation [181] . . . . . . . . . . . . . . . . . . . . . . . . 2.21 Results from static loading tests on 12.8 m pile ([61] after [14]) . . . 2.22 Extended modification factor Ff∗ for a. Chiibishi sand, b. Dogs Bay and c. Toyoura sand ([97]) . . . . . . . . . . . . . . . . . . . . . . . . 3.1. 3.2. 3.3. 3.4 3.5. a. Comparison of the limit surfaces Matsuoka-Nakai and Mohr Coulomb for different triaxial compression friction angles, φtc ; b. Geometrical representation of the invariants tanψ and cos3ϑ in principal stress space (after [158]) . . . . . . . . . . . . . . . . . . . . . . . . . a. Decrease of the maximum void ratio ei , the critical void ratio ec and the minimum void ratio ed with an increase of the mean effective pressure p b. Interaction between e, p and the density factor fd [22] Different intergranular strains δ related with different deformation histories. Only the recent part of the previous strain path (bold arrow) has an influence on δ. Current stress, void ratio and strain at current state * may be same for all three cases (after [120] ). . . . . . . . . . Evolution of intergranular stiffness M a. for χ = 6, b. for χ = 2; ρ varies from 0 to 1 [25]. . . . . . . . . . . . . . . . . . . . . . . . . . . Comparison of original (dashed lines) and modified (solid lines) reference ratios: a. change of e∗i0 , e∗c0 and e∗d0 by σv ; b. change of e∗i , e∗c and e∗d evolutions (after [129]). . . . . . . . . . . . . . . . . . . . . .. 22 23 24 24 25 25. 26. 28 29 30 32. 39. 40. 41 42. 44.

(24) List of Figures 3.6 3.7. Reference void ratio evolutions for primary loading - unloading - reloading cycles a. [129]; b. this study . . . . . . . . . . . . . . . . . . Comparison of Oedometer compression curves for reference hypoplastic model [158] and grain crushing models, [129] and improved crushing model (this study) . . . . . . . . . . . . . . . . . . . . . . .. Details on the Press-Replace modelling technique and progress of penetration of the pile . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2 Oedometer curves for different stress rate definitions and the PR method (after [153]) . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3 FE model for the volumetric analysis of a cone penetrating in an undrained clay: a. global view of the mesh configuration at the beginning of analysis; b. details on the mesh: green wedge indicating the embedded cone head, inclined slices prepared for the cone to be penetrated. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4 Surface heave near the cone at the end of 0.40 m penetration (max.uy = 6.67 · 10−3 m) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.5 Relation between Ir and Nc obtained from PR and other techniques (α = 1 full frictional/ rough, α = 0 frictionless/smooth contact). . . 4.6 Current plastic zones during steady state for soil rigidity indices: a) Ir = 150 and b) Ir = 300, full frictional contact; c) Variation of plastic zone with soil rigidity index based on simulation results of RITSS (after Lu et al. [106]) . . . . . . . . . . . . . . . . . . . . . . 4.7 Plastic zones around the penetrating cone simulated using explicit adaptive meshing with contact (after Walker & Yu [160] at different penetration depths (Ir = 100 and α = 1.0, full frictional contact). . . 4.8 Plastic zones around the penetrating cone simulated using PR technique at different penetration depths (Ir = 100 and α = 1.0, full frictional contact). . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.9 Penetration of a frictionless embedded conical pile. Normalised horizontal (left column) and vertical (right column) stresses (σxx /su and σyy /su , respectively) for a. zipper technique by prescribed displacements (Plaxis 2D) b. zipper technique by contact algorithm (Abaqus) c. PR technique (Plaxis 2D) . . . . . . . . . . . . . . . . . . . . . . . 4.10 Comparison of distribution (smoothness) of effective mean stresses (p ) of cases with high order (15-noded) elements (coarse mesh) and lower order (6-noded) elements (medium mesh) at 1m penetration level. 4.11 FE meshes of slice thickness of a. ts = 1 cm, b. ts = 2.5 cm, c. ts = 5 cm, and d. ts = 10 cm used in the comparison of step size and mesh density (Dark grey is the pile material). . . . . . . . . . . . . . 4.12 Penetration resistance curves of cases with varied step sizes for two different meshes uy = 2.5cm (blue lines) and uy = 5cm (red lines), namely. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. xxi. 47. 48. 4.1. 54 57. 59 59 60. 61. 62. 62. 65. 67. 68. 69.

(25) xxii. List of Figures. 4.13 Penetration resistance curves of cases with step sizes of a. uy = 5 cm, and b. uy = 10 cm for different slice thicknesses (ts = 2.5 cm, ts = 5 cm and ts = 10 cm) . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.14 Penetration resistance curves of cases with varied interface stiffness . 4.15 Normalised CPU times of all cases. . . . . . . . . . . . . . . . . . . . 4.16 a. Mobilised friction angle, φmob (first row) b. Void ratio, e (second row) distribution at penetration level, z= 1.00 m for Cases 2, 5, 7 & 10. 4.17 Assumed stress-strain curves for the interfaces a. at the extensions b. at the pile tip . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.18 Comparison of penetration curves for cases with peak (φpeak ) and critical (residual) interface strength parameter (φcritical ) . . . . . . . 4.19 Unloading issue evident from the load displacement curves (D = 0.40 - Id = 0.80) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.20 Stress distributions at previous phase (uy = 3.12 m) a. τ k−1 ; b σyy k−1 k−1 or σyy for horizontal interfaces) and c. σxx or σn for vertical interfaces to be activated in the current phase uy = 3.16 m . . . . . 4.21 Comparison of available interface strength and the shear stress at the beginning of step a. for horizontal interfaces at the pile tip, b. for vertical interfaces at the pile shaft . . . . . . . . . . . . . . . . . . . 4.22 Improvement of load displacement (FT - uy ) behaviour by using very strong interface extensions . . . . . . . . . . . . . . . . . . . . . . . . 5.1 5.2 5.3. 5.4. 5.5 5.6. 5.7 5.8 5.9. a. Sketch of the problem modelled b. General view of the FE model FE meshes prepared for PR analyses for cases with a. D = 0.30 m, b. D = 0.40 m . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . a. Illustration of the method for the calculation of the equivalent base force at the pile tip; b. Top view of one radian slice and the area of ith slice . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Total pile shaft, Fs , and base, Fb reaction curves obtained for piles having diameters and penetrating in sands with different soil densities, Id . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Load test curves for different pile diameters and total penetration lengths . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Comparison of results of the analysis using the PR technique with the design load test a. base resistance, b. shaft resistance curves suggested by the Dutch code (NEN 9997-2010) . . . . . . . . . . . . Sketch of the problem analysed and the description of 3D surface plots Idealised stiffness profiles of typical a. overconsolidated clay, b. sand, c. soft clay (after [111]) . . . . . . . . . . . . . . . . . . . . . . . . . Comparison of a. the effect of density; b. penetration length, and c. pile diameter on the distribution of normalised horizontal stresses, Rrr . In the leftmost column the 3D surface plots from a reversed angle of view, in the middle column the vertical cross-section at 0.01D from the shaft, and in the rightmost column the horizontal cross-sections at 0.5D below the tip are presented. . . . . . . . . . . . . . . . . . .. 69 71 72 73 73 74 75. 75. 76 77 82 83. 84. 85 86. 87 88 89. 90.

(26) List of Figures 5.10 Comparison of a. the effect of density ; b. penetration length, and c. pile diameter on the distribution of normalised vertical stresses, Rzz . In the leftmost column the 3D surface plots from a reversed angle of view, in the middle column the vertical cross-section at 0.01D from the shaft, and in the rightmost column the horizontal cross-sections at 0.5D below the tip are presented. . . . . . . . . . . . . . . . . . . 5.11 Comparison of a. the effect of density; b. penetration length, and c. pile diameter on the distribution of normalised tangential stresses, Rθθ . In the leftmost column the 3D surface plots from a reversed angle of view, in the middle column the vertical cross-section at 0.01D from the shaft, and in the rightmost column the horizontal cross-sections at 0.5D below the tip are presented. . . . . . . . . . . . . . . . . . . 5.12 Comparison of a. the effect of density; b. penetration length, and c. pile diameter on the distribution of normalised shear stresses, Rrz . In the leftmost column the 3D surface plots from a reversed angle of view, in the middle column the vertical cross-section at 0.01D from the shaft, and in the rightmost column the horizontal cross-sections at 0.5D below the tip are presented. . . . . . . . . . . . . . . . . . . 5.13 Comparison of a. the effect of density; b. penetration length, and c. pile diameter on the distribution of normalised void ratio, Re . In the leftmost column the 3D surface plots from a reversed angle of view, in the middle column the vertical cross-section at 0.01D from the shaft, and in the rightmost column the horizontal cross-sections at 0.5D below the tip are presented. . . . . . . . . . . . . . . . . . . 5.14 Determination of the equivalent critical state mean stress, pcs in hypoplasticity . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.15 Comparison of normalised a. base and b. shaft resistances during pile penetration together with trend lines obtained for Leighton Buzzard sand*[90]; and c. average void ratio evolution under pile base with ec reference curve (dashed line) of the hypoplastic model using Baskarp sand parameters. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.1 6.2 6.3 6.4. 6.5. Convex Hull of a PR discretisation (only sand region considered). . . Voronoi diagram and Delaunay tessellation (after [13]) . . . . . . . . Nearest Neighbour algorithm for a. 1D uniform, b. 2D uniform data, defined by stepwise discrete functions. . . . . . . . . . . . . . . . . . Nearest Neighbour algorithm for a scattered surface data represented by stepwise discrete functions based on the regions defined by Voronoi diagram; a. 2D representation, b. 3D representation . . . . . . . . . a. Voronoi diagram based on original data points 1,2,3,4 & 5; b. Second order Voronoi cells around the sample point X where the ratio of the enclosed areas Aabf, Aaehf, Aehd, Adcgh & Abcgf to Aabcde are used to determine the percent contribution from neighbouring cells of original data as weights [144]. . . . . . . . . . . . . . . . . . .. xxiii. 91. 92. 93. 94 94. 95 99 100 101. 101. 102.

(27) xxiv 6.6 6.7. 6.8. 6.9 6.10 6.11 6.12 6.13 6.14 6.15 6.16 6.17 6.18 6.19 6.20 6.21 6.22 6.23 6.24. 6.25. List of Figures The extended line of points and original mesh stress points (D = 30, L = 10D). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103 FE discretisation of a. PR, b) PR coarse (PR-C), and two extreme discretisation cases: c. finest mesh (F3) and d. coarsest mesh (C3) for case D = 0.30 m L = 10D . . . . . . . . . . . . . . . . . . . . . . 104 Comparison of normalised load displacement behaviour of different discretisations to which the results of the equalisation level of the PR analysis were imposed by interpolation. Lower curves are the load test results for K0 state, i.e. installation effects were not imposed. . 106 Comparison of a. imposed, and b. equalised interpolation distribution of σzz . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 107 Deviations from imposed (interpolated) state. . . . . . . . . . . . . . 108 Geometrical properties of the grid prepared to precondition the distribution of surface data for bivariate nonlinear regression calculations.113 Comparison of window size effect of the median filter on the σrz distribution of D = 0.30 m; L = 2.5D; Id = 0.40 . . . . . . . . . . . . . 115 Comparison of filtered data with PR data of σrz distribution of D = 0.30 m; L = 2.5D; Id = 0.40 a. σrr , b. σzz , c. σθθ , d. σrz ,and e. e . 116 Fitting results of normalised vertical stresses ψˆrr of cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . 117 Fitting results of normalised vertical stresses ψˆzz of cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . 118 Fitting results of normalised hoop stresses ψˆθθ of cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . 118 Fitting results of normalised shear stresses ψˆrz of cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . 119 Fitting results of normalised void ratios ψê of cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . 119 with PR data of cases: a. D = Comparison of the surface fit of σrr 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . 120 with PR data of cases: a. D = Comparison of the surface fit of σzz 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . 120 of cases: a. D = 0.30m; L = 10D; Comparison of the surface fit of σθθ Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . . . . . . . 121 with PR data of cases: a. D = Comparison of the surface fit of σrz 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . 121 Comparison of the surface fit of e with PR data of cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . 122 Comparison of numerical load test curves of the imposed fields using the fitting results with PR result and K0 state of cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . 122 Change of approximated field imposed (red), after equilibrium (soil, pink) and after pile and interfaces were activated (pile+NiL, green) a. radial and b. vertical directions of case D = 0.30m; L = 10D; Id = 0.60123.

(28) List of Figures 6.26 Distribution of radial unbalanced stresses for cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . 6.27 Distribution of vertical unbalanced stresses for cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . 6.28 Approximation result for the coefficient a11 and a31 (of ψˆrr ) in L/D, Id space. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.29 Approximation result for the coefficient b31 and r31 (of ψˆrr ) in L/D, Id space. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . distribution of generalised form with 6.30 Comparison of radial stress, σrr PR results of the cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . . . . . . . . . . . . . . . distribution of generalised form 6.31 Comparison of vertical stress, σzz with PR results of the cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . . . . . . . . . . . . . . . distribution of generalised form 6.32 Comparison of tangential stress, σθθ with PR results of the cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . . . . . . . . . . . . . . . 6.33 Comparison of shear stress, σrz distribution of generalised form with PR results of the cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . . . . . . . . . . . . . . . 6.34 Comparison of void ratio, e distribution of generalised form with PR results of the cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . . . . . . . . . . . . . . . . . 6.35 Comparison of radial unbalanced stresses for cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . 6.36 Distribution of vertical unbalanced stresses for cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . 6.37 Comparison of the numerical load test curves based on the equalised approximate installation effects with the state obtained using PR and the reference K0 state of cases: a. D = 0.30m; L = 10D; Id = 0.60, b. D = 0.40m; L = 10D; Id = 0.80 . . . . . . . . . . . . . . . . . . . 7.1 7.2. 7.3. xxv. 124 124 127 127. 128. 129. 129. 130. 130 131 131. 132. FE mesh used for the verification cases a. D = 0.30 m and b. D = 0.40 m . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 134 Comparison of the numerical load test curves of the PR results (interpolated on the reference mesh), the equalised approximate installation effects and the reference K0 results of the verification cases a. D = 0.30 m; Id = 70; and b. D = 0.40 m; Id = 50, both having a pile length, L = 6D . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 135 Comparison of the numerical load test curves of the PR results (interpolated on the reference mesh), the equalised approximate installation effects and the reference K0 results of the verification cases a. D = 0.30 m; Id = 70; and b. D = 0.40m; Id = 50, both having a pile length, L = 8D. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 135.

(29) xxvi 7.4 7.5. 7.6. 7.7. 7.8 7.9 8.1. List of Figures a. Physical model test setup [56] (all units are in mm; acceleration level, Ng = 35), and b. FE model of the validation case . . . . . . . Comparison of the numerical load test curves with the centrifuge test result [56] based on the equalised approximate installation effects with the state obtained using formulas and the reference K0 state of the dense case, e0 = 0.637; Id = 0.77. . . . . . . . . . . . . . . . . . . . . Comparison of the numerical load test curves with the centrifuge test result [56] based on the equalised approximate installation effects with the state obtained using formulas and the reference K0 state of the medium dense case, e0 = 0.709; Id = 0.58. . . . . . . . . . . . . . . . Comparison of the numerical load test curves with the centrifuge test result [56] based on the equalised approximate installation effects with the state obtained using formulas and the reference K0 state of the loose case, e0 = 0.783; Id = 0.38. . . . . . . . . . . . . . . . . . . . . Comparison of total pile capacities . . . . . . . . . . . . . . . . . . . Comparison of pile a. base and b. shaft capacities . . . . . . . . . .. 137. 138. 138. 139 141 141. Comparison of the numerical load test curves using formulas employed with using original stiffness (hs ) and high stiffness (100 · hs ) for the medium dense case, e0 = 0.709;Id = 0.58. . . . . . . . . . . . . . . . 151.

(30) List of Tables 3.1 3.2 4.1 4.2 4.3 4.4 4.5 4.6 5.1 5.2. 6.1 6.2 6.3 7.1 7.2 7.3 7.4. Hypoplastic soil model parameters for Baskarp sand [8] . . . . . . . Estimated grain crushing parameters [129] for the modified grain crushing model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 47. Parameters and results of the cases analysed for volume preservation check (c = 50 kPa) . . . . . . . . . . . . . . . . . . . . . . . . . . . . Parameters and results of the cases analysed for volume checking control Hypoplastic soil model parameters for Baskarp sand (after [8]) . . . Summary of analysis for the investigation of step size . . . . . . . . . Summary of analysis for the investigation of mesh density . . . . . . Summary of cases analysed in parametric study on the interface stiffness. 59 63 66 68 68 71. Hypoplastic soil model parameters for Baskarp sand (after [8]) . . . Summary of variations analysed for the investigation of pile diameter (D), penetration length (L) and density (Id ) effects on the installation field . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. 47. 82. 82. Geometrical properties of the regions defined for the interpolation grid 113 Summary of regression analysis of a)D = 0.30m; L = 10D: Id = 0.60 and b)D = 0.40m; L = 10D: Id = 0.80 . . . . . . . . . . . . . . . . . 117 Summary of regression analysis on the generalisation of the normalised radial stresses . . . . . . . . . . . . . . . . . . . . . . . . . . . . 128 Summary of improvement levels of the verification cases at uy = 0.10D136 Summary of improvement levels of the validation cases at uy = 0.10D 137 Summary of pile total capacity prediction accuracies of proposed functions and empirical methods . . . . . . . . . . . . . . . . . . . . . . . 142 Summary of pile base and shaft capacity prediction accuracies of proposed functions and empirical methods . . . . . . . . . . . . . . . . . 142. xxvii.

(31) xxviii. List of Tables.

(32) List of Symbols Latin c Cu D Deq d50 d60 , d10 e e d , ec , ei ed0 , ec0 , ei0 ref Eoed , Eoed f Ff ∗ Ft , Fs , F b g G Gi , Gs h hs Id Ir K k S , kN K0 K0N C L m mR , mT n Ng Nq p. cohesion coefficient of uniformity pile diameter radius of the area of influence for density change effects mean grain size sieve opening size through which 60% and 10% of the particles pass void ratio minimum, critical, and maximum void ratio at current stress reference minimum, critical, and maximum void ratio at low stress oedometric stiffness at current and reference stress levels load vector modification factor considering compressibility of the sand pile head, shaft, and base reaction gravitational acceleration shear modulus shear modulus of interface, and soil distance from pile tip granulate hardness density index rigidity index global stiffness matrix interface shear and normal stiffness at rest earth pressure coefficient at rest earth pressure coefficient for normally consolidated soil pile (penetration) length exponent of stress dependent stiffness hypoplastic model stiffness modification factors number of piles within the influence area or exponent of granulate hardness centrifuge acceleration ratio bearing capacity factor effective mean stress xxix.

(33) xxx pa pref p0 P1 , P2 , P3 q b , qc , qs qb , max Qu r R Rs s S SD smax SP T − N te , ts u ur uy v vc r, vf. List of Symbols atmospheric pressure (100 kPa) reference mean stress effective mean stress at geostatic stress state contact forces on a particle P1 > P2 > P3 base, cone and shaft resistance pile base resistance at maximum settlement level ultimate load capacity radial distance or normalised centre to centre distance b/w adjacent piles pile radius or particle roundness normalised shaft resistance settlement particle sphericity normalised displacement maximum settlement the standard penetration resistance number element, and slice thickness displacement displacement normal to pile shaft vertical prescribed displacement applied at pile head specific volume offset of crushing and end of crushing lines. Greek α α p , αq , βp , βq β. δ δ Δe Δσr ε˙ H η φ φc φtc ψ ψˆrr , ψˆzz , ψˆθθ , ψˆrz , ψê λ0 , λcr , λf. exponent controlling volumetric behaviour in hypoplastic law or interface strength reduction factor modification factors for reference void ratios normalised shaft resistance or exponent controlling volumetric behaviour in hypoplastic law frictional coefficient of pile-soil interface intergranular strain tensor change in void ratio change in normal effective stress on pile shaft during loading due to change in soil volume strain rate Henky strain improvement factor internal friction angle of the soil critical friction angle triaxial compression friction angle state parameter model functions of Cartesian stress and void ratio fields slope of the critical, crushing and end of crushing state.

(34) xxxi. λ σ ˚ σ σm , σrc σrf σt σv0 σxx σyy σ 1 , σ3 τmax , τres τrz τs. lines load scaling factor Cauchy stress Jaumann stress rate mean normal stress radial effective stress acting on pile shaft at failure and after installation tensile stress in a particle geostatic effective vertical stress horizontal stress vertical stress major and minor principal stress peak and residual shear strength shear stress at pile shaft, at axisymmetric conditions shear strength of the pile shaft.

(35) xxxii. List of Symbols.

(36) Chapter 1 Introduction 1.1. Background. One of the most traditional methods to support structures resting on soft soils is the use of piles. They generally work by transferring the loads to deeper soil layers, which can provide enough bearing capacity when mobilised. This type of foundations has been commonly used around the world. In the western part of The Netherlands this type of foundation is often used due to the typical soil profile which consists of a thick (10-20 m) soft soil layer underlain by a stiff bearing stratum composed mostly of quartz sand. Piles need to be strong and durable. In the past, wooden piles were used sometimes decay in time and hence lose their carrying capacity. Quite common tilted Dutch houses are clear indication of the significance of this issue or overloading due to the neglected downdrag. In current practice, piles made up of pre-cast reinforced concrete or steel are used. Piles can be installed in different ways. The most common way to install the prefabricated piles is the use of an impact hammer. In this driving method, the pile is displaced by each drop or blow of the ram until the desired penetration level is achieved. Another alternative is the vibratory driving technique where the pile is forced to penetrate by a heavy vibratory head on top. The vibrations degradate the strength of (or liquefy if saturated) the surrounding soil and due to the heavy weight of the vibrator the pile penetrates. Recent developments have allowed the piles to be jacked resulting in reduced nuisance and higher capacity and stiffness as compared to the hammering and vibratory driving techniques ([55], [76]). The bearing capacity of a pile depends on the soil properties and the stress state. The behaviour of granular material is governed by the packing of the grains and the contact stresses between. The stress and density dependent behaviour determines the soil response. The effects of the installation operation are transmitted to the soil through the interaction of sand grains and the pile, resulting in an altered soil state and properties. The installation effects depend on the applied installation technique. The change of stress and density as well as change of the physical properties (e.g. grain size dis1.

(37) 2. 1. Introduction. tribution) of the soil in the vicinity of the pile installed can be quite different for different installation methods. The resulting confining stresses of a dynamic installation technique (hammering or vibratory driving) are lower than those of the jacking technique. Depending on the density, dilation would be expected near the pile shaft during jacking (e.g. [38], [56]). On the other hand, dynamic installation methods can cause compaction or loosening near the pile depending on the method and state of the soil. For example impact driving, and for the loose and medium dense sands vibratory driving, causes compaction ([83]), while for the dense sand vibratory driving causes loosening in the vicinity of the pile. The numerical simulations of [112] indicated compaction around the pile for all dynamic installation methods and all soil densities. Two important consequences of the installation effects that should be considered in the design are: The influence on the performance of the pile (e.g. load–displacement behaviour, capacity) in its service life ([102], [56]). For example, if the density of the soil as well as the confining stress around the pile is high, a higher bearing capacity will be mobilised with less displacement. The influence on the neighbourhood. For example, dynamic installation techniques have often nuisance (e.g. noise, vibrations in the neighbourhood) issues ([171], [118]) in populated areas. Furthermore, the technique can cause damage on the neighbouring structures [172]. The induced vibrations, depending on the soil type, might amplify the vibrations and cause compaction of foundation strata. That can cause structural damage to the neighbouring structures, mostly due to induced settlements of the foundation soil. In the static penetration case (e.g. jacking) the neighbouring underground structures (e.g. prefabricated pile, sheet pile wall) may be influenced by the resulting displaced soil and the increased stress level ([76]).. A more realistic behaviour and therefore an improved design would be achieved by considering the installation effects in the analyses than performing the analysis considering a geostatic stress state around the pile(s) modelled. In the current practice, the installation effects are taken into account by some empirical design methods in order to estimate the bearing capacity of foundation piles [123]. Several field and model tests performed to investigate the influence of pile installation on the bearing capacity, have led to only an evolution in empirical models to estimate the bearing capacity of displacement piles [87], [125]. Coop et al. [46] attempted to take the soil state around the pile into account by measuring both the contact stresses on the pile as well as displacements of the sand around the pile in model tests, but did not directly measure density changes in the zone influenced by the pile installation. Recent attempts to investigate the change in the soil state were also limited to either the measurement domain (generally close to the pile) or resolution as well as the variables (e.g. displacement, strain, stress, density) that can be quantified [164], [56], [103]. However, the behaviour of piles during installation as well as the interaction with the surrounding soil and the change of soil properties during.

(38) 1.1. Background. 3. installation are still not well known. This information is essential, not only to make better predictions of the pile bearing capacity and its behaviour in the soil under different loading conditions, but also to predict the (side) effects of pile installation on the neighbourhood. The increase in computational power allowed improved numerical techniques like the finite element method (FEM); however installation effects of displacement piles have not been incorporated in the numerical models for engineering purposes. As a result of discarding the installation effects there will be a large difference between the FE predictions and the real behaviour (performance) of the pile [34]. Therefore it is important to model the installation effects of displacement piles numerically so that the load displacement behaviour of the pile during its service life as well as the secondary effects on neighbouring structures would be predicted more accurately. For a proper analysis of the pile installation and its consequent effects, all important aspects should be included in the numerical model. These can be summarised as: Large deformation effects, which require geometry update and use of objective stress-strain measures (e.g. Truesdell, Green-Naghdi, Jaumann). When modelling the penetration there are local large deformations that cause stress concentrations and distortion of the mesh. Unless properly modelled the analysis will result in quite a different state than that would be expected. For example geometry should be updated in order to account for penetration and obtain more realistic displacement, strain and therefore stress field. Furthermore, the stress (and conjugate strain) measures should be objective (e.g. Jaumann rate of stress) so that the rigid body rotations would not cause any stress or strain increase [24]. Additional volumetric and shear deformation effects in the objective stress measure such as Hill and Truesdell rate of stress, respectively, would be more appropriate. However, to obtain a realistic response, these objective stress rates should be checked for certain stress–strain paths (like 1-D compression, [153]). Pore pressure effects, which have to be considered when the loading is fast enough to cause undrained behaviour of the soil. If the soil is fully saturated the pore pressure generation and dissipation should also be modelled. Inertia effects, which depend on the rate of loading, should also be considered for a proper modelling. The intrinsic properties of the material, the coupled system (e.g. saturated dense sand) as well as the rate of loading determine the response of the system. The displacements, velocities and accelerations mobilize corresponding components of the system. As a conclusion, inertia effects should be considered in the analysis of dynamic installation techniques. Use of an appropriate constitutive model is quite important to mimic common characteristics of the soil considered. As most common bearing strata is quartz sand in the Netherlands, the model used should have a state dependent behaviour (e.g. Critical State Models) in which the soil stiffness and strength is dictated by current state. Furthermore, at high stress levels the sand grains.

(39) 4. 1. Introduction can be crushed. As a result, the material properties will change. Therefore, the effect of grain crushing should be also considered in the constitutive model.. Nevertheless, simplifications are inevitable. For pile jacking, which is considered as a quasi-static penetration, the inertia effects can be discarded and drained conditions can be assumed (no excess pore pressure effects). Therefore in the resulting FE model, the level of complexity will be reduced.. 1.2. Scope and Limitations. This study is the numerical part of a larger project in which the installation effects of displacement piles are investigated both experimentally and numerically. The objective of this numerical study is to investigate the installation effects of pile jacking in sand in a numerical framework. Since the jacking operation can be considered as quasi-static loading, the dynamic effects are not considered in the analyses. For the same reason, drained conditions are assumed such that the pore pressures are only taken into account as a reduction of total stress (giving the effective stresses). As a common tool used in engineering practice, FE modelling is selected as the numerical method. In order to model pile jacking using the FEM in a standard FE code, some additional aspects have to be considered to account for the large deformation effects. In this perspective possibilities to simulate installation effects as well as the installation process were investigated and a simplified technique (the Press-Replace technique) is proposed. After a sensitivity study, the method is employed to investigate the installation effects around a jacked pile for different sand densities. In the view of the results obtained from FE simulations, the possibilities of incorporating the installation fields were investigated. Furthermore, describing the installation effects in a functional form, which allows flexibility for imposing the installation field around a so-called wished-in-place pile to model it as a jacked pile, was studied. A novel technique is introduced to describe the installation effects of a jacked pile without simulating the penetration process. This way a jacked pile can be modelled wished-in-place and corresponding installation effects (i.e. stress, density and stiffness change) can be imposed. By the proposed method, there will be an enormous gain in term of computational effort and time in the analyses of displacement piles.. 1.3. Outline. The thesis consists of 8 chapters. First current practice on investigation and design methods of the pile installation and consequent effects are explained (Chapter 2). The important features of the constitutive model used in the analyses and the grain crushing modifications are presented (Chapter 3). The Press-Replace technique, which is developed for standard FE algorithm, is explained (Chapter 4). The method is employed to investigate geometric and material effects on the installation field.

(40) 1.3. Outline. 5. Chapter 5. The possibilities of the representation of the installation effects and their incorporation in a standard FE code is presented in Chapter 6. The method for incorporation of the installation effects is validated by a centrifuge model test in Chapter 7. Finally, in Chapter 8 the findings of this study are summarised and some recommendations for future research are given..

(41) 6. 1. Introduction.