Foundations Design of a New Hospital Building in KECSKEMÈT

FOUNDATION DESIGN OF A NEW HOSPITAL BUILDING IN KECSKEMÉT LUÍS MIGUEL CUNHA E COSTA Dissertação submetida para satisfação parcial dos requisitos do grau de MESTRE EM ENGENHARIA CIVIL — ESPECIALIZAÇÃO EM GEOTECNIA Orientador: Professor Doutor António Joaquim Pereira Viana da Fonseca Coorientador: Professor Doutor András Mahler JULHO DE 2013 MESTRADO INTEGRADO EM ENGENHARIA CIVIL 2012/2013 DEPARTAMENTO DE ENGENHARIA CIVIL Tel. +351-22-508 1901 Fax +351-22-508 1446  [email protected] Editado por FACULDADE DE ENGENHARIA DA UNIVERSIDADE DO PORTO Rua Dr. Roberto Frias 4200-465 PORTO Portugal Tel. +351-22-508 1400 Fax +351-22-508 1440  [email protected]  http://www.fe.up.pt Reproduções parciais deste documento serão autorizadas na condição que seja mencionado o Autor e feita referência a Mestrado Integrado em Engenharia Civil - 2012/2013 - Departamento de Engenharia Civil, Faculdade de Engenharia da Universidade do Porto, Porto, Portugal, 2013. As opiniões e informações incluídas neste documento representam unicamente o ponto de vista do respetivo Autor, não podendo o Editor aceitar qualquer responsabilidade legal ou outra em relação a erros ou omissões que possam existir. Este documento foi produzido a partir de versão eletrónica fornecida pelo respetivo Autor. Foundation Design of a New Hospital Building in Kecskemét Ao meu avô Bernardino Foundation Design of a New Hospital Building in Kecskemét i ACKNOWLEDGMENTS The author would like to express his appreciation and gratitude to some people who contributed decisively to the final result of the current dissertation. To Professor António Viana da Fonseca, firstly, for having made possible the current work in Budapest and after for all the help and valuable suggestions during the work, despite the physical distance. To Professor András Mahler, for accepting the co-supervision proposal, for suggesting the thesis subject and for all the recommendations and help provided during my stay in Budapest. To Structural Engineer Gábor Schnierer, for the help provided during the model implementation in AxisVM. To my Friends, for the friendship during all these years, for the bohemian moments, for the good conversations and for the care. We are Bon Vivants! To my Brother Gonçalo, for the great friendship and for the unthinkable moments of humor shared. To my girlfriend Catarina, for all the motivation, support and specially for the selfless love. Finally and most importantly to my Parents, for believing in me and for the unconditional love and support. Foundation Design of a New Hospital Building in Kecskemét ii Foundation Design of a New Hospital Building in Kecskemét iii ABSTRACT The goal of the current work was the design of a new hospital building located in Kecskemét, Hungary, focusing essentially geotechnical aspects. In this way, the calculation procedure took into account both the ultimate limit state and the serviceability limit state. Initially, it was treated the CPTU data collected in the site, which permitted the correct definition of soil stratigraphy, resistance and deformability properties and subsequently the accurate assessment of the most effective foundation system. After opting for a deep foundation solution, an ultimate state calculation was performed in order to determine the capacity of a single pile, clearly identifying the influence of each of its components, base resistance and shaft resistance. The obtained results led to a possible piled raft system. The stress-strain analysis was carried out by a computer program based on the finite element method, AxisVM, and compared with a simplified hand calculation method. As usual in deformation problems analysis, associated to less severe consequences for the structure, a serviceability limit state is defined. The reinforcement of both piles and raft was performed under the ultimate limit state assuring moderate concrete pressures to avoid cracking. AxisVM was used to perform the referred calculations. Finally, a drainage system solution was proposed considering the initial high level of the water table. To run this calculation the Flow Mode of the program Plaxis 2D was selected and the solution found was compared with an empirical hand calculation method. KEYWORDS: Bearing Capacity, CPTU test, Deep Foundations, Drainage, Piles, Raft, Settlement. Foundation Design of a New Hospital Building in Kecskemét iv Foundation Design of a New Hospital Building in Kecskemét v RESUMO O objetivo do presente trabalho foi o dimensionamento de um edifício de um novo hospital localizado em Kecskemét, Hungria, focado essencialmente nos aspectos da Engenharia Geotécnica. Neste sentido, no processo de cálculo foram tidos em conta os estados limite último e de serviço. Inicialmente, foi tratada a informação relativa aos ensaios CPTU efetuados in situ que permitiram a correta definição da estratigrafia do solo, das características de resistência e deformabilidade e consequentemente a avaliação mais eficaz do sistema de fundações. Depois da opção por um sistema de fundações profundas, foi determinada a capacidade de carga de uma única estaca em estado limite último, evidenciando claramente a influência de cada uma das suas componentes, resistência de ponta e resistência lateral. Os resultados obtidos permitiram obter uma possível solução de estacas. A análise tensão-deformação foi elaborada através de um programa baseado no Método dos Elementos Finitos, AxisVM e comparado com o método simplificado de cálculo manual. Tal como é habitual na análise de problemas de deformação, o dimensionamento é condicionando pelo estado limite de serviço. O cálculo da armadura das estacas e do maciço de encabeçamento obedeceu ao estado limite último assegurando de tal modo tensões moderadas no betão evitando a fendilhação deste. O programa AxisVM foi de novo utilizado para este cálculo. Por fim, foi proposta uma solução de drenagem tendo em conta a posição inicial elevada do nível freático. Este cálculo foi efetuado no programa Plaxis 2D através da opção Flow e mais uma vez a solução obtida foi confrontada com um método mais simplificado, desta feita de natureza empírica. PALAVRAS-CHAVE: Assentamento, Capacidade de Carga, Drenagem, Ensaio CPTU, Estacas, Fundações profundas, Maciço de Encabeçamento. Foundation Design of a New Hospital Building in Kecskemét xii Foundation Design of a New Hospital Building in Kecskemét xiii LIST OF FIGURES Figure 2.1 – Detail of the construction area (Petik and Partners Engineering Service Ltd).................... 3 Figure 2.2 – Overview of the constructed building .................................................................................. 4 Figure 2.3 – Geological Map of Kecskemét, Bács-Kiskun County, scale 1:100000 (Gyalog, 2005) ...... 4 Figure 2.4 – Cross section A-A (Petik and Partners Engineering Service Ltd) ....................................... 5 Figure 2.5 – Cross section B-B (Petik and Partners Engineering Service Ltd) ....................................... 6 Figure 2.6 – Relative density and consistency index (FUGRO Consult KFT) ........................................ 6 Figure 2.7 – CPT1 reported data (FUGRO Consult KFT) ....................................................................... 7 Figure 2.8 – CPT1 reported data within the adopted classification chart (FUGRO Consult KFT) .......... 7 Figure 2.9 – CPT2 reported data (FUGRO Consult KFT) ....................................................................... 8 Figure 2.10 – CPT2 reported data within the adopted classification chart (FUGRO Consult KFT) ........ 8 Figure 2.11 – CPT3 reported data (FUGRO Consult KFT) ..................................................................... 9 Figure 2.12 – CPT3 reported data within the adopted classification chart (FUGRO Consult KFT) ........ 9 Figure 2.13 – Detail of a typical CPT probe (Robertson, 2012) ............................................................ 12 Figure 2.14 – Pore pressure filter locations (Lunne et al.,1997) ........................................................... 12 Figure 2.15 – First chart based on CPT test (Begemann, 1965) .......................................................... 14 Figure 2.16 – Simplified soil type classification chart (Robertson and Campanella, 1983) .................. 14 Figure 2.17 – Soil behaviour type classification chart based on CPTU data (Lunne et al., 1997) ........ 15 Figure 2.18 – Normalized soil behaviour type classification chart based on CPTU data (Lunne et al., 1997) ..................................................................................................................................................... 16 Figure 2.19 – Normalized soil behaviour type classification chart (Robertson, 2012) .......................... 17 Figure 2.20 – Unit weight correlation chart with CPT data (Mayne, 2007) ........................................... 20 Figure 2.21 – Young’s Modulus vs. depth ............................................................................................. 24 Figure 2.22 – Linear trend lines adjusting the Young’s Modulus evolution ........................................... 25 Figure 3.1 – Components of a pile bearing capacity ............................................................................. 29 Figure 3.2 – Components of the pile end resistance in EC7 approach................................................. 30 Figure 3.3 – Load settlement curves: a) Observed curve based on a load test; b) Theoretical curve . 34 Figure 3.4 – Example of results obtained in CEFEUP/ISC’2 Experimental Site (Viana da Fonseca and Santos, 2008) modelled by Fernandes (2011) ...................................................................................... 35 Figure 3.5 – Theoretical unit shaft resistance curve ............................................................................. 35 Figure 3.6 – Theoretical unit base resistance curve ............................................................................. 36 Figure 3.7 – Illustration of the vertical strain influence factor calculation (adapted from www.rocscience.com) ........................................................................................................................... 37 Figure 3.8 – Young’s modulus vs. depth (Mayne, 2001) ...................................................................... 40 Figure 3.9 – Example of interaction factor method (Poulos, 2006) ....................................................... 41 Figure 4.1 – Detail structural plan view (Bohn Deep Foundation Ltd) .................................................. 43 Figure 4.2 – On the left, a photo of a CFA pile and on the right, a photo of a OMEGA pile ................. 44 Figure 4.3 – Execution sequence of CFA piles (adapted from www.junttan.com) ................................ 45 Figure 4.4 – Execution sequence of a general screw pile (adapted from www.junttan.com) ............... 46 Foundation Design of a New Hospital Building in Kecskemét xiv Figure 4.5 – Relation between the skin friction and the effective cone resistance (Mahler, 2007) ....... 50 Figure 4.6 – Equivalent pl depending on the soil type (Bustamante and Gianeselli, 1996) .................. 53 Figure 4.7 – Qi Abacus correlating Ménard Limit Pressure (pLM) with skin friction (qs) (Bustamante et al., 2009) ................................................................................................................................................ 57 Figure 4.8 – Design loads applied to the columns (Bohn Deep Foundation Ltd) .................................. 61 Figure 4.9 – Organizational chart detailing the calculation procedure .................................................. 63 Figure 4.10 – Initial OMEGA piles solution ............................................................................................ 64 Figure 4.11 – Overlapping piles detail ................................................................................................... 65 Figure 4.12 – Final OMEGA pile solution .............................................................................................. 65 Figure 5.1 – “Nodal support” option in “Elements” menu (adapted from AxisVM) ................................ 68 Figure 5.2– Load-settlement curve for 0.4 m diameter pile ................................................................... 69 Figure 5.3 – Load-settlement curve for 0.6 m diameter pile .................................................................. 69 Figure 5.4 – “Surface support” option in “Elements” menu (adapted from AxisVM) ............................. 70 Figure 5.5 – Iε depth evolution for L/B=2.56 .......................................................................................... 71 Figure 5.6 – 3D perspective of the defined model (adapted from AxisVM) ........................................... 73 Figure 5.7 – Case study 1 Isosurface 2D (adapted from AxisVM) ........................................................ 73 Figure 5.8 – Case study 2 Isosurface 2D (adapted from AxisVM) ........................................................ 74 Figure 5.9 – Case study 3 Isosurface 2D (adapted from AxisVM) ........................................................ 74 Figure 5.10 – Case study 4 Isosurface 2D (adapted from AxisVM) ...................................................... 74 Figure 5.11 – Evolution of the interaction factor with the spacing between piles .................................. 77 Figure 6.1 – 3D perspective of the raft foundation (adapted from AxisVM) .......................................... 82 Figure 6.2 – Cover to reinforcement (adapted from AxisVM) ................................................................ 83 Figure 6.3 – Plan xOy (adapted from AxisVM) ...................................................................................... 83 Figure 6.4 – Actual reinforcement window on AxisVM software (adapted from AxisVM) ..................... 84 Figure 6.5 – Reinforcement difference xt (adapted from AxisVM) ........................................................ 85 Figure 6.6 – Reinforcement difference xb (adapted from AxisVM) ....................................................... 85 Figure 6.7 – Reinforcement difference yt (adapted from AxisVM) ........................................................ 85 Figure 6.8 – Reinforcement difference yb (adapted from AxisVM) ....................................................... 86 Figure 7.1 – Water conditions in the construction site ........................................................................... 87 Figure 7.2 – Adopted stratigraphy for the groundwater calculations ..................................................... 88 Figure 7.3 – Modelled structure in Plaxis 2D (adapted from Plaxis 2D) ................................................ 89 Figure 7.4 – Generated mesh in Plaxis 2D (adapted from Plaxis 2D) .................................................. 89 Figure 7.5 – Boundary conditions window in Plaxis 2D (adapted from Plaxis 2D)................................ 90 Figure 7.6 – Evolution of the water table position (adapted from Plaxis 2D) ........................................ 90 Figure 7.7 – Lowering level obtained with 7.5 length trenches (adapted from Plaxis 2D) .................... 91 Figure 7.8 – Left trench |q| values (m3/day/m) (adapted from Plaxis 2D) ............................................. 91 Figure 7.9 – Right trench |q| values (m3/day/m) (adapted from Plaxis 2D) ........................................... 91 Figure 7.10 – Dewatering for trenchworks (Puller, 2003) ...................................................................... 92 Foundation Design of a New Hospital Building in Kecskemét xv LIST OF TABLES Table 2.1 – Soil Behaviour Type zones defined in the Unified Approach (Robertson, 2012) ............... 18 Table 2.2 – CPT1 results for the soil behaviour type (SBT) .................................................................. 18 Table 2.3 – CPT2 results for the soil behaviour type (SBT) .................................................................. 19 Table 2.4 – CPT3 results for the soil behaviour type (SBT) .................................................................. 19 Table 2.5 – CPT1 average qc and fs values .......................................................................................... 19 Table 2.6 – CPT2 average qc and fs values .......................................................................................... 19 Table 2.7 – CPT3 average qc and fs values .......................................................................................... 19 Table 2.8 – Estimated fs values ............................................................................................................. 20 Table 2.9 – Estimated unit weight values .............................................................................................. 21 Table 2.10 – Proposed αE values .......................................................................................................... 22 Table 2.11 – Adopted αE values ............................................................................................................ 23 Table 2.12 – Estimation of the E based on CPT1 results ..................................................................... 23 Table 2.13 – Estimation of the E based on CPT2 results ..................................................................... 23 Table 2.14 – Estimation of the E based on CPT3 results ..................................................................... 23 Table 2.15 – Reliability ratings regarding design with CPTU data (adapted from Lunne et al., 1997) . 26 Table 3.1 – αp values (adapted from Lakatos, 2011) ............................................................................ 31 Table 3.2 – αs;g values (adapted from Lakatos, 2011) .......................................................................... 32 Table 3.3 – αs;c values (adapted from Lakatos, 2011) .......................................................................... 32 Table 3.4 – Vertical strain influence factor expressions (Matos Fernandes, 2011) .............................. 38 Table 4.1 – Hypothetic qc data .............................................................................................................. 48 Table 4.2 – Pile end bearing capacity (pmax;base ) according to EC7 ...................................................... 49 Table 4.3 – End bearing force (Rb;cal) according to EC7 ....................................................................... 49 Table 4.4 – CPT1 qs values for CFA and OMEGA piles ....................................................................... 50 Table 4.5 – CPT2 qs values for CFA and OMEGA piles ....................................................................... 50 Table 4.6 – CPT3 qs values for CFA and OMEGA piles ....................................................................... 51 Table 4.7 – Side friction force (Rs;cal) according to EC7 ........................................................................ 51 Table 4.8 – Average and minimum values of compressive resistance ................................................. 51 Table 4.9 – Design total resistance of the piles after the correlation factors ........................................ 51 Table 4.10 – qce values for the different diameters considered............................................................. 52 Table 4.11 – Pile end bearing capacity (qb) according to LCPC Method .............................................. 52 Table 4.12 – Pile end bearing force (Rb;cal) according to LCPC Method ............................................... 52 Table 4.13 – β correlation factor and qs;max based directly on CPT data (Viana da Fonseca et al., 2012) ............................................................................................................................................................... 53 Table 4.14 – CPT1 equivalent pl values ................................................................................................ 54 Table 4.15 – CPT2 equivalent pl values ................................................................................................ 54 Table 4.16 – CPT3 equivalent pl values ................................................................................................ 54 Table 4.17 – Description of the 418 analysed piles (Bustamante et al., 2009) ..................................... 55 Table 4.18 – Pile type Curves (Qi) associated to a certain soil type (Bustamante et al., 2009) ........... 56 Foundation Design of a New Hospital Building in Kecskemét xvi Table 4.19 – Respective curves Qi adopted.......................................................................................... 56 Table 4.20 – CPT1 qsi values for CFA and OMEGA piles ..................................................................... 57 Table 4.21 – CPT2 qsi values for CFA and OMEGA piles ..................................................................... 57 Table 4.22 – CPT3 qsi values for CFA and OMEGA piles ..................................................................... 58 Table 4.23 – Side friction force (Rs;cal) according to LCPC Method ...................................................... 58 Table 4.24 – Average and minimum values of compressive resistance ............................................... 58 Table 4.25 – LCPC design total resistance of the piles after the correlation factors ............................ 58 Table 4.26 – Comparison between the Rc;d obtained by the two methods ........................................... 59 Table 4.27 – Relation between the pile type and the respective fclim and k1 (DTU 13.2, 1992) ............ 59 Table 4.28 – k2 values ........................................................................................................................... 60 Table 4.29 – fc* values ........................................................................................................................... 60 Table 4.30 – Concrete compressive strength limits .............................................................................. 60 Table 4.31 – Rc;DTU values ..................................................................................................................... 61 Table 4.32 – Comparative design total resistance of the piles .............................................................. 62 Table 4.33 – Design total resistance of the piles ................................................................................... 62 Table 4.34 – Average price of both CFA and OMEGA piles ................................................................. 64 Table 4.35 – Equivalent force application point of the two piles............................................................ 65 Table 5.1 – Settlements associated to resistance mobilization ............................................................. 68 Table 5.2 – Load-settlement curve data for 0.4 m diameter pile ........................................................... 68 Table 5.3 – Load-settlement curve data for 0.6 m diameter pile ........................................................... 68 Table 5.4 – Maximum characteristic loads ............................................................................................ 69 Table 5.5 – Pile flexibility for each diameter .......................................................................................... 70 Table 5.6 – Equivalent rectangular raft dimensions .............................................................................. 70 Table 5.7 – Raft settlement ................................................................................................................... 71 Table 5.8 – Modulus of subgrade reaction ............................................................................................ 72 Table 5.9 – Stiffness parameters input .................................................................................................. 72 Table 5.10 – Maximum and average settlements ................................................................................. 75 Table 5.11 – Maximum and average differential settlements ................................................................ 75 Table 5.12 – Maximum and average angular distortions ...................................................................... 75 Table 5.13 – A and B combinations ....................................................................................................... 76 Table 5.14 – αij values ........................................................................................................................... 76 Table 5.15 – Maximum and average settlements ................................................................................. 77 Table 5.16 – Maximum and average differential settlements ................................................................ 78 Table 5.17 – Maximum and average angular distortions ...................................................................... 78 Table 5.18 – Maximum and average settlement comparison................................................................ 78 Table 5.19 – Maximum and average differential settlement comparison .............................................. 79 Table 5.20 – Maximum and average angular distortion comparison .................................................... 79 Table 6.1 – Minimum pile reinforcement area ....................................................................................... 82 Table 6.2 – Required reinforcement for the raft design ......................................................................... 84 Foundation Design of a New Hospital Building in Kecskemét xvii Table 7.1 – Right trench |q| values (m3/day/m) ..................................................................................... 92 Table 7.2 – Hand calculation results ..................................................................................................... 93 Table 7.3 – Comparison between both methods .................................................................................. 93 Foundation Design of a New Hospital Building in Kecskemét xviii Foundation Design of a New Hospital Building in Kecskemét xix NOMENCLATURE Latin Characters A – coefficient to determine the pile interaction factor (α) Ab – pile cross-sectional area (m2) Ac – projected area of the cone (m2) Ac – pile cross-sectional concrete area (m2) AN – projected area associated to cone inner diameter (m2) AS – area of friction sleeve (m2) As – longitudinal reinforcement area (m2) As;bpmin – minimum longitudinal reinforcement area (m2) Asi – shaft area associated to soil layer “i” (m2) AT – projected area associated to cone outer diameter, equal to Ac (m2) B – smaller width of a rectangular foundation (m) B – coefficient to determine the pile interaction factor (α) Bq – pore pressure parameter C – coefficient to determine the pile interaction factor (α) Cs – linear correction factor D – cone outer diameter (m) D – pile diameter (m) D – coefficient to determine the pile interaction factor (α) Db – diameter of pile base (m) Deq – equivalent pile diameter (m) d – cone inner diameter (m) eZ – displacement in vertical (z) direction E – Young’s modulus (MPa) Eb – soil Young’s modulus below foundation base (kPa, MPa) Ep – pile modulus (kPa, MPa) EsL – Young’s Modulus at the pile tip (kPa, MPa) Esm – soil Young’s modulus at mid-depth of pile shaft (kPa, MPa) Fc;d – design load applied to a certain column, at the ultimate limit state (kN) Fk;max – maximum characteric load applied to a pile (kN) Foundation Design of a New Hospital Building in Kecskemét xx fcj – characteristic compressive cylinder strength of concrete at j days (MPa) fclim – limit strength of concrete depending on pile foundation type (MPa) fc28 – compressive cylinder strength of concrete at 28 days (MPa) fc* – concrete compressive strength (MPa) Fs – total force acting on friction sleeve (kN) fs – unit sleeve friction resistance (kPa, MPa) Gs – specific gravity of the soil H – height of the static water table (m) Hc – total height of clay strata (m) hD – maximum residual head (m) hj – thickness of stratum “j” (m) Hs – total height of sand strata (m) h0 – height of the water table in wells (m) Ic – soil behaviour type index Iε – vertical strain influence factor Iεc – continues foundation Iε Iεp – peak of Iε Iεs – axisymmetric foundation Iε Iρ – displacement influence factor Kc – pile tip bearing factor k – modulus of subgrade reaction (kN/m3) k – soil coefficient of permeability (m/s) kc – clay coefficient of permeability (m/s) keq – equivalent soil coefficient of permeability (m/s) ks – sand coefficient of permeability (m/s) kx – horizontal soil coefficient of permeability (m/s) ky – vertical soil coefficient of permeability (m/s) kp – pile flexibility (kN/m) ks – soil modulus of subgrade reaction (kN/m3) k1 – coefficient that takes into account the execution method of the pile k2 – coefficient that takes into account the difficulty related to the concrete casting of the pile L – pile length (m) L – greater width of a rectangular foundation (m) Foundation Design of a New Hospital Building in Kecskemét xxi M – constrained deformation modulus (MPa) n – number of piles per column p – raft distributed load (kN/m2) Pav – average load on a pile within the group (kN/m) pa – atmospheric stress (kPa, MPa) pl – Ménard Limit Pressure (kPa, MPa) pLM – Ménard Limit Pressure (kPa, MPa) PT – unit load applied on pile I (kN) pmax;base – maximum base resistance (kPa, MPa) Q – flow (m3/s/m, m3/day/m, L/day/m) Qavg – average flow (m3/s/m, m3/day/m, L/day/m) Qb – pile end bearing capacity (kN) qb – pile unit end resistance (kPa, MPa) Qc – total force acting on the cone (kN) qc – measured cone resistance (MPa) qce – cone resistance around the base (kPa, MPa) qc;I;mean – component of pmax;base formula (kPa, MPa) qc;i;avg – the layer “i” average qc (kPa) qc;II;mean – component of pmax;base formula (kPa, MPa) qc;III;mean – component of pmax;base formula (kPa, MPa) qE – effective cone resistance (kPa, MPa) Qs – pile shaft friction capacity (kN) qs – pile shaft friction resistance (kPa, MPa) qsi – pile unit side friction (kPa, MPa) qsi;c – pile unit side friction in cohesive stratum (kPa) qsi;g – pile unit side friction in granular stratum (kPa) qs;max – upper limit of shaft friction (kPa, MPa) Qult – ultimate pile axial capacity (kN) |q| – absolute mode of flow (m3/s/m, m3/day/m, L/day/m) |q|avg – average absolute mode of flow (m3/s/m, m3/day/m, L/day/m) Rb – pile end bearing resistance (kN) Rb;cal – pile base resistance, calculated from ground test results, at the ultimate limit state (kN) (Rb;cal)mean – average of Rb;cal in the tests performed (kN) Foundation Design of a New Hospital Building in Kecskemét 2 The assessment of the piled raft settlements was also carefully discussed and analysed. Finally, considering the site water conditions, a dewatering system in the excavation pit was also designed. 1.3. ORGANIZATION OF THE THESIS The present dissertation is divided in eight chapters and one appendix. The defined structure was considered adequate to cover all the proposed subjects, starting with more general chapters (2 and 3), where initial considerations are addressed and the methodologies are described, followed by the calculation and the respective result presented in chapters 4, 5, 6 and 7. In Chapter 1, a brief introduction of the present work is made, including the scope and objectives and the organization of the thesis. In Chapter 2, the geological and geotechnical framework of the construction site is defined, including the foundations type choice, the adopted soil stratigraphy and the estimation of soil strength and deformability parameters. Chapter 3 presents all the methodologies used to perform the calculation of both the pile single bearing capacity and piled raft settlements. In Chapter 4, the results regarding pile single capacity are summarized, as well as the choice between the two execution methods studied for piles (CFA and OMEGA), leading to a possible piled foundation system. Chapter 5 presents the settlements estimation, where a numerical method is confronted with an empirical methodology. In Chapter 6 all the reinforcement calculations are included, for the both the piles and the slab. The same numerical program referred in the previous chapter was used to perform the reinforcement slab calculation. The pile reinforcement was calculated according to Eurocode 2. Chapter 7 presents a possible dewatering solution for the executed excavation. Again, a numerical calculation was performed and compared with an empirical method. Finally, in Chapter 8, the final observations are made, regarding the main conclusions of the present work. In Appendix A is presented the characteristic and design loads applied on the piles as well as a plan view of the design loads over the columns. Foundation Design of a New Hospital Building in Kecskemét 3 2 GEOLOGICAL AND GEOTECHNICAL FRAMEWORK OF THE CONSTRUCTION SITE 2.1. PROJECT DETAILS The object of study of this work is the “A Building” foundations design, where area is highlighted in Figure 2.1 with a red contour. Figure 2.1 – Detail of the construction area (Petik and Partners Engineering Service Ltd) Foundation Design of a New Hospital Building in Kecskemét 4 The hospital’s working platform is at the 112.400 mBf level from an initial ground level ranging from 116.29 mBf (drilling 43RNF) to 116.82 mBf, meaning an excavation varying from 3.89 m to 4.42 m. The inclination of the slopes is 3:2. As it was referred previously, the objective was the geotechnical foundation design of a new hospital located in Kecskemét, constructed in 2009 (Figure 2.2). Figure 2.2 – Overview of the constructed building 2.2. GEOLOGY The hospital’s construction area, represented on the geological map (Figure 2.3), is located in formations deposited during different phases of the Quaternary Period, being some of them relatively recent. The initial letter of the abbreviators in the legend of Figure 2.3, ‘Q’, represents the referred geological period. Figure 2.3 – Geological Map of Kecskemét, Bács-Kiskun County, scale 1:100000 (Gyalog, 2005) Foundation Design of a New Hospital Building in Kecskemét 5 The sediments of the Quaternary form a thin superficial layer of approximately 10 m, being composed essentially by coarse grained deposits. The older subsequent stratum consists of compact clayey layers and alluvial sediments, which are loose and formed by the mixture of sand and clay. Underlain a thousand meters thick Upper Pannonian lacustrine sequence, arises Fluvial Quaternary deposits. 2.3. GEOTECHNICAL OVERVIEW AND CHARACTERIZATION During the survey, several boreholes were executed that allowed the definition of soil stratigraphy associated to the working area. In the plan view shown in Figure 2.1 are represented the respective locations of the seven drilling holes, that allowed the definition of two cross sections, A-A and B-B. The depth of drillings are twelve meters (1F, 2F and 3F) and sixteen meters (I.DSz, II.DSz and 44RNF), with dynamic probing heavy (DPH) tests – not used in the further geotechnical interpretation – performed alongside the boreholes I.DSz and II.DSz. The cross section A-A (Figure 2.4) shows an embankment of an anthropic clayey deposit 1.5 m thick, followed by a slightly organic silt-clay stratum with two meters thickness. Around the four meters depth, the drillings performed distinguished two different soil types, on the left the predominance of sandy silt ground and on the right silty sand. Figure 2.4 – Cross section A-A (Petik and Partners Engineering Service Ltd) The cross section B-B (Figure 2.5) presents a very identical geotechnical sequence when compared with cross section A-A. The first two layers have the same characteristics and the last sequence, starting also around four meters depth, is a silty sand stratum. Foundation Design of a New Hospital Building in Kecskemét 6 Figure 2.5 – Cross section B-B (Petik and Partners Engineering Service Ltd) In both cases, the determined phreatic level position is 1.55 m depth. The evaluation of the groundwater levels is made in depth in all the boreholes, through the installation of a piezometer connected to the surface, where the results can be read when the sensor reaches the water level. In order to determine the soil behaviour, regarding resistance, stiffness and deformability, 3 CTP/CPTU tests were performed, one the most renowned characterization ground tests in the current geotechnical practice, further outlined in this work (Subchapter 2.5). The logs (unit sleeve friction resistance, fs, cone resistance, qc, friction ratio, Rf) of the three CPT tests are shown in Figure 2.7, Figure 2.9 and Figure 2.11, alongside with the respective soil behaviour type classification chart (Figure 2.8, Figure 2.10 and Figure 2.12). The colours in the Rf log correspond to one of the zones identified in the classification chart, whereas the colours in the qc log must associated with either relative density (Dr) or consistency index (Ic) in the case of sands or clays, respectively (Figure 2.6). Figure 2.6 – Relative density and consistency index (FUGRO Consult KFT) Foundation Design of a New Hospital Building in Kecskemét 7 Figure 2.7 – CPT1 reported data (FUGRO Consult KFT) Figure 2.8 – CPT1 reported data within the adopted classification chart (FUGRO Consult KFT) Foundation Design of a New Hospital Building in Kecskemét 8 Figure 2.9 – CPT2 reported data (FUGRO Consult KFT) Figure 2.10 – CPT2 reported data within the adopted classification chart (FUGRO Consult KFT) Foundation Design of a New Hospital Building in Kecskemét 9 Figure 2.11 – CPT3 reported data (FUGRO Consult KFT) Figure 2.12 – CPT3 reported data within the adopted classification chart (FUGRO Consult KFT) Foundation Design of a New Hospital Building in Kecskemét 10 2.4. ASSESSMENT OF THE FOUNDATIONS TYPE In foundation design, the consideration of a shallow foundation as a first approach is the most common solution, taking into account the lower costs involved and the simpler implantation complexity when compared to the adoption a deep foundation system. The two basic types of shallow foundations are isolated footings, working independently, and mat foundations, i.e., a concrete slab assuring a redistribution of the loads. The deep foundations can be associated to a basement slab (frequently called raft) by integrating a group of piles into a system, called piled raft. The choice of the foundations type is based mainly in the following aspects:  Soil stiffness;  Soil strength;  Structural loads;  Foundation dimensions In the specific case of this work, considering the reduced average qc values immediately below the working ground level (112.400 mBf) and the dimensions the building object of this design (around 1000 m2), a piled raft foundation system was adopted. Some advantages of this option are the greater settlement control due to the interaction between the piles and the slab and a more reduced stress level in the concrete slab, consequence of an uniform load distribution. Another conclusion that can also be taken regards to pile length, a primary input of any ultimate pile capacity method. Again recalling the CPT data on Figure 2.7, Figure 2.9 and Figure 2.11, the qc values are relatively high between 15 m and 18 m depth below the initial ground surface, which appeared to be an adequate position to set the piles tip. Therefore, the pile tip position was chosen to be 16 m (100.53 mBf) depth meaning a 11.87 m pile length, based on the defined working ground level (112.40 mBf). In practice, the considered pile length can be said to be 12 m. 2.5. THE CPT TEST FOR THE DESIGN OF DEEP FOUNDATIONS 2.5.1. GENERAL DESCRIPTION OF THE CPT TEST The Cone Penetration Test is one of the most significant in situ tests used in the common geotechnical practice (Lunne et al., 1997). Considering that the first developments of the mechanical CPT started in the first half of the 20th century, it is logically a very well-founded test, used by many researchers for ground characterization. The CPT test answers to central questions of the soil characterization being the one the most important tests used in the geotechnical design practice. The main advantages of the CPT are the continuous profiling, the reliable data registered and the strong theoretical background developed through the years which permits reliable correlations. On the other hand, this test requires skilled operators, relatively high investment and also does not provide soil sample. Although the Standard Penetration Test (SPT) remains the most widely in situ test used in the world (Matos Fernandes, 2011), due to the its simplicity in operation and low cost, the CPT proved to be a much more reliable test, since it provides the measurement of three physical quantities, cone tip resistance, qc, sleeve friction, fs, and pore pressure, u2, the latter register giving a new designation to the test, CPTU as it is going to be referred further. One of the biggest limitations of the SPT is the measurement of only one parameter (in the case, the number of blows, N) for the purpose of assuring a full characterization of the soil in study. The CPT has filled that lack by using different physical parameters which at the same time might converge to confirm a certain soil behaviour. Foundation Design of a New Hospital Building in Kecskemét 11 The first cone penetrometer tests were made in 1932, as referred above, with a cone tip attached to the steel rod, with 10 cm2 of projected area and an apex angle of 60º and were performed by the Dutch engineer Pieter Barentsen in Netherlands. In 1953, Begemann improved considerably the Dutch cone penetration test with the measurement of the local sleeve friction, possible with the introduction of an “adhesion jacket” behind the cone. Thus, it was possible to register not only the cone tip resistance but also the sleeve friction, which leaded to the first soil type classification based on CPT (Begemann, 1965). An electric cone was developed in 1965 by Fugro with the dimensions and shape that were adopted as basis for the modern cones and also for the normalized test procedures (e.g. ASTM). The main advantages of the electric penetration relative to the mechanical part were the elimination of the incorrect readings of the sleeve friction, the continuous soil profiling possible the continuous penetration ratio and also the more accurate measurements provided by an electric device. Regarding the evolution of the equipment, the year of 1974 brought eventually the most remarkable introduction in the history of the test. The cones started featuring a filter ring (piezocones), where it was possible to measure the pore water pressure. This innovation consequently led to an effective stress analysis, instead of the erroneous total stress interpretation hitherto used. The procedure of the test consists on the penetration of a cone placed at the end of a series of rods that is pushed into the ground at a constant rate and provides simultaneously continuous measurements of the cone resistance, sleeve friction and even, when it is used the “piezocone”, the pore pressure. The measurement of the pore pressure is nowadays very common in the geotechnical practice considering the advantages that it brings regarding the full characterization of the soil. The use of the piezocone instead of the original electric CPT implied also the introduction of a correction in the measured values for the cone resistance, being used a parameter called qt. The ring filter located between the end of the friction sleeve and the cone creates an unequal area effect (Campanella et al., 1982), generating water pressures in descendent direction in the upper part of the cone. This pressure added to the measured cone resistance, qc, correspond to the total resistance of the soil, qt. Thus, (2.1) where, The second part of equation 2.1 will be greater the smaller the inner diameter, d, is. Currently, the equipment used to perform the piezocone has values for the parameter a between 0.70 and 0.85. From equation 2.1, it can be also concluded that the unequal area effect could be very significant in the total cone tip resistance, mainly when the parameter assumes the lower limit of the range. It still must be understood that this correction is only pertinent for cohesive soils, as result of the excess pore water pressure induced by the penetration. In granular soils, the value of qc can be considered equal to qt, regarding the drained behaviour that dominates in this type of materials. A similar correction must also be applied to the sleeve friction values when the cross sectional areas on the top and the bottom of the Foundation Design of a New Hospital Building in Kecskemét 18 Table 2.1 – Soil Behaviour Type zones defined in the Unified Approach (Robertson, 2012) Zone Soil Behaviour Type Ic 1 Sensitive, fine grained N/A 2 Organic soils – clay > 3.60 3 Clays – silty clay to clay 2.95 – 3.60 4 Silt mixtures – clayey silt to silty clay 2.60 – 2.95 5 Sand mixtures – silty sand to sandy silt 2.05 – 2.60 6 Sands – clean sand to silty sand 1.31 – 2.05 7 Gravelly sand to dense sand < 1.31 8 Very stiff sand to clayey sand* N/A 9 Very stiff, fine grained* N/A * Heavily overconsolidated or cemented The zones can be included once more in two main groups:  From 1 to 4, cohesive soils;  From 5 to 9, granular soils. 2.5.3. ADOPTED SOIL STRATIFICATION In the scope of this work, the CPT interpretation data provided by the responsible company was made through the SBT charts developed by Robertson et al. (1986). Using the friction ratio (Rf) log in Figure 2.7, Figure 2.9 and Figure 2.11 and associating its colours to the ones in the respective classification charts (Figure 2.8, Figure 2.10 and Figure 2.12), it was possible to define two main soil behaviour types:  Granular;  Cohesive. This distinction is very common and important, since almost every CPT based design method suggests different coefficients whether the soil is sand or clay. The adopted soil stratification is presented in Table 2.2, Table 2.3 and Table 2.4. Table 2.2 – CPT1 results for the soil behaviour type (SBT) SBT Layer hi (m) hf (m) H (m) Granular 1 112.40 112.00 0.40 Cohesive 2 112.00 110.70 1.30 Granular 3 110.70 106.50 4.20 Cohesive 4 106.50 105.50 1.00 Granular 5 105.50 100.53 4.97 Foundation Design of a New Hospital Building in Kecskemét 19 Table 2.3 – CPT2 results for the soil behaviour type (SBT) SBT Layer hi (m) hf (m) H (m) Granular 1 112.40 112.00 0.40 Cohesive 2 112.00 111.00 1.00 Granular 3 111.00 106.50 4.50 Cohesive 4 106.50 103.00 3.50 Granular 5 103.00 100.53 2.47 Table 2.4 – CPT3 results for the soil behaviour type (SBT) CPT3 Layer hi (m) hf (m) H (m) Granular 1 112.40 111.80 0.60 Cohesive 2 111.80 111.00 0.80 Granular 3 111.00 100.53 10.47 Associated to each of the layers defined previously, corresponds an average value of the cone tip resistance (qc) and the sleeve friction (fs) as follows in Table 2.5, Table 2.6 and Table 2.7: Table 2.5 – CPT1 average qc and fs values Layer qc,mean (MPa) fs,mean (kPa) 1 26.34 218.2 2 3.92 39.8 3 9.89 94.9 4 1.52 24.3 5 10.00 76.3 Table 2.6 – CPT2 average qc and fs values Layer qc,mean (MPa) fs,mean (kPa) 1 24.44 257.0 2 8.35 97.6 3 15.39 159.7 4 3.16 67.0 5 21.64 187.9 Table 2.7 – CPT3 average qc and fs values Layer qc,mean (MPa) fs,mean (kPa) 1 7.83 89.6 2 1.92 18.9 3 13.39 116.3 Foundation Design of a New Hospital Building in Kecskemét 20 2.6. ESTIMATION OF RESISTANCE AND DEFORMABILITY PARAMETERS 2.6.1. UNIT WEIGHT OF DEFINED STRATUM The estimative of the unit weight (γ) was made based on the Figure 2.20, which correlates the parameter with the sleeve friction (fs) and the specific gravity of the solid, Gs. Figure 2.20 – Unit weight correlation chart with CPT data (Mayne, 2007) According to Table 2.5, Table 2.6 and Table 2.7 a weighted average of fs was defined, for both granular and cohesive soil types, taking into account the thickness of the layers. For example, the granular sleeve friction average value associated to the CPT1 was calculated as follows: (2.9) Table 2.8 summarizes the weighted fs values for both soil behaviour types in each of the CPT tests performed. The last column includes the adopted values. Table 2.8 – Estimated fs values SBT fs,CPT1 (kPa) fs,CPT2 (kPa) fs,CPT3 (kPa) fs,average (kPa) Granular 90.4 174.4 114.9 126.6 Cohesive 33.1 73.8 18.9 41.9 Using the equation written in the Figure 2.20, it is possible to estimate the soil unit weight, as follows: Foundation Design of a New Hospital Building in Kecskemét 21 (2.10) The value adopted for the Gs parameter was 2.6 for both soil types, leading to the respective unit weights presented in the Table 2.9: Table 2.9 – Estimated unit weight values SBT fs,average (kPa) γ (kN/m3) Granular 126.6 18.0 Cohesive 41.9 16.7 2.6.2. ADOPTED VALUE FOR THE POISSON’S RATIO Since there was no data regarding the estimative of the Poisson’s Ratio (ν) in the scope of the geotechnical investigation, there were considered the average values concerning the both types of soils:  Granular soils: ν=0.3;  Cohesive soils: ν=0.2-0.3, considering a drained behaviour. 2.6.3. ESTIMATION OF YOUNG’S MODULUS The most common correlation used to determine the Young’s Modulus (E) via CPT test is expressed in equation 2.17. (2.11) where, α is the modulus factor. The value of alpha (α) is, in this case, a coefficient related to the Young’s modulus (E), being usually indicated in index the letter “E”. The equation 2.12 was proposed by Robertson (2009), depending on the soil behaviour type index, Ic. [ ] (2.12) Considering the two dominant behaviours aforementioned, a value for Ic was defined for both granular and cohesive materials, based on the ranges defined in Table 2.1. For granular soils, and using the classification charts on Figure 2.8, Figure 2.10 and Figure 2.12, it is possible to identify a concentration of points around the zone 8 and 9, meaning a soil behaviour type so-called silty sand to sand. Hence, correlating this classification with the data in Table 2.1, a value of 1.8 was defined for this material, corresponding to a silty sand. Foundation Design of a New Hospital Building in Kecskemét 22 For cohesive soils, the definition of a single soil behaviour type is not as simple as in the previous case, fact caused by the relative dispersion of points between the zones 3, 4, 5 and 6 (Figure 2.8, Figure 2.10 and Figure 2.12). Nevertheless, with the exception of the CPT2 results where there are a significant presence of marks in zone 3, in the other two tests the points belong mostly to border regions, 5 and 6. Hence, the cohesive soil behaviour type varies largely from silty clay to clayey silt, being defined for the SBT index the value of 3.0. In Table 2.10, the Young’s modulus cone factor (αE) values determined through equation 2.12 are summarized: Table 2.10 – Proposed αE values SBT Ic αE Granular 1.8 7.02 Cohesive 3.0 32.07 The obtained value for the granular soils is consistent with the density of this sand (dense to very dense) based on the relative density, Dr, plotted alongside with qc in each of the CPTU tests (Figure 2.8, Figure 2.10 and Figure 2.12). On the other hand, the calculated value for fine-grained soils exceeds largely the recommendations of several authors (Senneset, 1992; Kulhawy and Mayne, 1990), who indicate as upper limits for the constrained modulus cone factor (αM) of 15 the first, and 14 for the last two. In linear elasticity, the constrained modulus M is linked to the Young’s modulus E via: (2.13) Considering the Poisson's ratio adopted range ν=0.2-0.3, characteristic for cohesive soils with drained behavior, then: Conservatively adopting ν=0.3 and αM=14, it can be written: It should be noted that the proposed formula presented on equation 2.12 is associated to the net cone resistance (qt-σvo). However, this nuance was neglected having regarding the low influence (reducing the E value in 1.39%) of the total overburden stress when compared with qc averages. In accordance with the above, the adopted αE values are resumed in the Table 2.11. Foundation Design of a New Hospital Building in Kecskemét 23 Table 2.11 – Adopted αE values SBT αE Granular 7.02 Cohesive 10 Table 2.12, Table 2.13 and Table 2.14 resume the average Young’s modulus in each layer. Table 2.12 – Estimation of the E based on CPT1 results CPT1 Layer H (m) qc,mean (MPa) E (MPa) Granular 1 0.40 26.34 184.8 Cohesive 2 1.30 3.92 39.2 Granular 3 4.20 9.89 69.4 Cohesive 4 1.00 1.52 15.2 Granular 5 4.97 10.00 70.1 Table 2.13 – Estimation of the E based on CPT2 results CPT2 Layer H (m) qc,mean (MPa) E (MPa) Granular 1 0.40 24.44 171.5 Cohesive 2 1.00 8.35 83.5 Granular 3 4.50 15.39 108.0 Cohesive 4 3.50 3.16 31.6 Granular 5 2.47 21.64 151.8 Table 2.14 – Estimation of the E based on CPT3 results CPT3 Layer H (m) qc,mean (MPa) E (MPa) Granular 1 0.60 7.83 54.9 Cohesive 2 0.80 1.92 19.2 Granular 3 10.47 13.39 93.9 After setting these coefficients, it is important to define a tendency or in other words to assign a law for the Young’s modulus evolution in depth. Thus, it was considered a linear evolution of the modulus starting in the surface with a nonzero value, considering the obtained results through the ground tests. The obtained values for the Young’s modulus, associated to the midpoint of each layer, are represented in Figure 2.21. Foundation Design of a New Hospital Building in Kecskemét 24 Figure 2.21 – Young’s Modulus vs. depth As can be seen, the evolution with depth shows to be very inconstant as result of the measured qc variability. Despite that, it shall be mentioned that some of those results are influencing too much considering their global importance for the geotechnical behaviour of the foundation, such as:  The extremely high qc values in the first layer are justified by the presence of gravels or even rocks that must be neglected considering its low thickness (0.20 m, 0.20 m, 0.30 m for CPT1, CPT2, CPT2, respectively);  In CPT1 test, around the 6 m depth, the specific average qc value (1.52 MPa) for the cohesive stratum affects excessively the linear results, considering it is only 1 m thick. Thus, the referred data should not be considered in the weighing. In relation to CPT2, the low average value in the cohesive layer (4) at 7.65 m depth must also be removed from the trend line calculation, adjusting a posteriori the equation by introducing a factor that takes into account the relevance of the mentioned stratum. Figure 2.22 shows the simplified Young’s modules curves and the respective linear correlation, an equation of the type y=mx+b. 0 2 4 6 8 10 12 050 100 150 200 z (m) E (MPa) CPT1 CPT2 CPT3 Foundation Design of a New Hospital Building in Kecskemét 25 Figure 2.22 – Linear trend lines adjusting the Young’s Modulus evolution The equations related to each ground test are as follows: (2.24) (2.35) (2.16) Regarding the equation 2.15, it should also be applied a correction due to the exclusion of the layer 4 from the previous calculations. The factor is equal to: where, Hlayer 4 is the thickness of layer 4; Htotal is the total depth of analysed CTP data. 0 2 4 6 8 10 12 050 100 150 200 z (m) E (MPa) CPT1 CPT2 CPT3 Linear (CPT1) Linear (CPT2) Linear (CPT3) Foundation Design of a New Hospital Building in Kecskemét 26 Hence, the equation 2.15 is now written as follows: (2.17) The adopted equation is the average of the previous three, as follows: (2.18) 2.7. APPLICATION OF THE CPT/CPTU RESULTS IN DEEP FOUNDATIONS DESIGN One of the major applications of the test was, since its beginning, and continues to be the determination of the bearing capacity of piles. The continuous soil profiling in depth allows the pile surrounding ground definition. However, the prediction of the pile capacity is not easily obtained since it depends on different factors, such as:  Type of the pile;  Installation method, related to the previous one;  Type of the soil crossed. The empirical CPT/CPTU design methods developed through the years tend to integrate all these effects, but the difficulty in quantifying them requires the introduction of factors of safety, through the reduction of the calculated values. All the design calculations were made following the Eurocode 7 (EC7) regulations, that preconizes the use of resistance and correlation factors, being the first ones applied to determine the characteristic value and followed by the calculation of the design values. This issue will be resumed in the next chapter. Regarding the adequacy of the CPT/CPTU data in design problems, Lunne et al. (1997) compiled the gained experience to develop the Table 2.15: Table 2.15 – Reliability ratings regarding design with CPTU data (adapted from Lunne et al., 1997) Soil Type Pile Design Bearing Capacity Sand 1-2 1-2 Clay 1-2 1-2 Intermediate soils 1-2 2-3 *Reliability ratings: 1High; 2High to moderate; 3Moderate; 4Moderate to Low; 5-Low According to Table 2.15, it can be confirmed the good reliability provided by the CPT/CPTU data in geotechnical design, specifically in deep foundations design. Moreover, Robertson (2012) stated that the CPT test is the most accurate in pile design. Foundation Design of a New Hospital Building in Kecskemét 27 3 METHODS OF ANALYSIS FOR SINGLE PILE CAPACITY AND STRESS-STRAIN BEHAVIOUR OF THE PILED RAFT 3.1. INTRODUCTION In geotechnical design project, the prediction of the resistance and deformability of the soil are the main goal. Hence, the calculation approaches must be carefully chosen, based essentially on its theoretical premises consistency, suitability to the case study and applicability in common practice. Considering the high uncertainty that surrounds geotechnical problems, the design approaches must be verified based on comparisons with other methods, avoiding unsafe solutions and framing the range of results. Thus, there were considered two static approaches for the single pile capacity prediction, both hand calculation methods. For the deformation analysis of the piled raft two approaches of different nature were considered, the first using a structural program based on finite element method (FEM) and the second based on an geotechnical hand calculation method. During the chapter the adopted methodologies are described in a general way, which will then be adapted to the specific case study. 3.2. CALCULATION OF THE BEARING CAPACITY OF A SINGLE PILE The value of the bearing capacity of a single pile (Qult) comprises two components:  End bearing load (point resistance), Qb (kN);  Side friction load (shaft resistance), Qs (kN). Thus, (3.1) The equation 3.1 can also be written in terms of resistance (R), as follows: ∑ (3.2) Foundation Design of a New Hospital Building in Kecskemét 34 3.3. STRESS-STRAIN BEHAVIOUR ANALYSIS The deformation analysis continues to be one the hardest issues to solve in the scope of the geotechnical project regarding the high non-linearity of the soil behaviour. Specifically, the soil stiffness, perhaps the most important parameter involved in the serviceability of the structures, tends to have a rather inconstant evolution with depth (2.6.3) being at the same time impossible to reflect all this variability. Thus, the pile group settlements assessment is achieved assuming some simplifications, in any case conservative, of the behaviour of not only the soil but also the piled raft and its interaction. As defined in Eurocode 7 (CEN, 2004), “a limiting value for a particular deformation is the value at which a serviceability limit state, such as unacceptable cracking, is deemed to occur in the supported structure”. 3.3.1. THEORETICAL LOAD-SETTLEMENT CURVES Considering the difficulties, mainly logistics, in relation to the execution of a considerable and representative number of pile load tests in the construction site, several projects assume idealized loadsettlement curves (see Figure 3.3) based on registers obtained in instrumented piles, namely in static load tests. Figure 3.3 – Load settlement curves: a) Observed curve based on a load test; b) Theoretical curve The theoretical curves are constructed by overlaying the shaft resistance, so called t-z curves and the base resistance, named as q-z curves, if both are expressed in stress along depth (z). Through the observation of results of load tests performed in CFA piles (as the one plotted in Figure 3.4, in an experimental site in residual soils), it is possible to establish a relation between the developed settlement at the ultimate base or shaft resistance with pile diameter. Foundation Design of a New Hospital Building in Kecskemét 35 Figure 3.4 – Example of results obtained in CEFEUP/ISC’2 Experimental Site (Viana da Fonseca and Santos, 2008) modelled by Fernandes (2011) A common and simplified method is to assume that the mobilization of each of the components occurs for a certain deformation level, expressed as function of the pile diameter. Obviously, the generated settlement depends on the pile diameter, as referred above, but also on the execution method. Hence, shaft resistance is generally fully mobilized for values around 1.5% to 2.0% (Figure 3.5) of the pile diameter (D1), developing a minor settlement and not depending significantly on the piles execution method. Figure 3.5 – Theoretical unit shaft resistance curve Foundation Design of a New Hospital Building in Kecskemét 36 On the other hand, the mobilization of the base resistance (Figure 3.6) requires a larger settlement being in this case more dependent on the pile type, with the values ranging from 5% (displacement piles) to 10% (replacement piles) of the pile diameter (D2). Figure 3.6 – Theoretical unit base resistance curve Thus, the pile stiffness is essentially given by the portion of the base resistance considering the early mobilization of the shaft resistance. 3.3.2. ANALYSIS OF RAFT SETTLEMENTS A raft can be compared to a footing of larger dimensions and its study can be done similarly. Thus, the settlement estimative of a raft can be done following the same assumptions adopted for the footings, where the elastic solutions are still used and valid unlike what happens in deep foundations considering this king of analysis. Notwithstanding, the empirical and semi-empirical methods are also accepted and well-recognized in geotechnical design practice mainly for the conservative solutions. The Schmertmann Method (1970) is an example of a semi-empirical method that derives from the theoretical elastic solutions and it was intended to narrow certain deviations observed between the referred solutions and real cases. The method is one of the most renowned in the scope of the settlement analysis of a shallow foundation regardless its remote introduction. A brief description of the method is done below. The general expression of the immediate settlement proposed by Schmertmann (1970) is presented in the equation 3.13: ∫( ) (3.13) where, Cs is a dimensionless linear correction factor calculated according to equation 3.14; Foundation Design of a New Hospital Building in Kecskemét 37 Δqs is the net foundation pressure; Iε is the vertical strain influence factor explained in Figure 3.7; E is the Young’s Modulus. ( ) (3.14) where, σ'vb is the effective vertical stress at foundation level. The equation 3.14 reflects the effect of the foundation embedment, logically favourable for the global settlement reducing it and a very significant enhancement introduced by the method. It must be referred that one of the causes for the deviation in the results of the elastic solution is related to nonconsideration of Cs factor. Actually, the elastic solutions do not consider the influence of the effective vertical stress increase with the depth, which leads to increase of the Young’s Modulus too. The Figure 3.7 describes graphically the evolution with depth of the vertical strain influence factor, where the vertical strain grows until reaching a peak (Iεp) depending on the stress level. Thereafter the influence factor starts to decrease, lowering until zero. The correspondent ordinate represents the vertical strain influence depth. The factor Iε is calculated according to Table 3.4 Figure 3.7 – Illustration of the vertical strain influence factor calculation (adapted from www.rocscience.com) Foundation Design of a New Hospital Building in Kecskemét 38 Table 3.4 – Vertical strain influence factor expressions (Matos Fernandes, 2011) In practice, the integral presented generally in equation 3.13 is replaced by a sum of n sublayers in which the soil is divided (equation 3.15): ∑ (3.15) 3.3.3. ESTIMATION OF PILE GROUP SETTLEMENTS The analysis of pile group settlements must focus the pile-soil interaction since each pile behaves itself not independently but rather interacting with the surrounding ones. In this way, the stress-strain behavior should assess not only the individual pile deformation due to the load application over its axis but also the effect that the group will induce in the global settlement. There are numerous approaches regarding the estimation of pile group settlements amongst them the well-recognized Interaction Factor Method proposed by Poulos and Davis (1980) which is described below. The global settlement (wi) of the pile i within a group of n piles is given by the equation 3.16: ∑( ) (3.16) where, Pav is the average load on a pile within the group; S1 is the settlement of a single pile under unit load; αij is the interaction factor for pile i due to any other pile (j) within the group, corresponding to the spacing sij between piles i and j. Analysing the equation 3.16, some considerations must be done regarding the determination of s1 and αij, since the value of the first component (Pav) is consequence of direct application of the project serviceability load. The settlement of a single pile under unit load (s1) corresponds to the pile flexibility (FL-1) and, assuming a continuous elastic behaviour of the soil, it can be determined as follows in the equation 3.17, as proposed by Poulos (1987): Foundation Shape Depth Iε expression Axisymmetric (square or circular) 0≤ z ≤B/2 B/2 ≤ z ≤ 2B Iε = 0.1+(z/B).(2.Iεp-0.2) Iε = 0.667Iεp.(2-z/B) Continues 10≤L/B 0≤ z ≤B B ≤ z ≤ 4B Iε = 0.2+(z/B).(Iεp-0.2) Iε = 0.333Iεp.(4-z/B) 1<L/B<10 Iε = Iεs+0.111(IεcIεs) Iεs is the Iε for an axisymmetric foundation Iεc is the Iε for an continues foundation Foundation Design of a New Hospital Building in Kecskemét 39 (3.17) where, PT is the unit load applied on pile I; Iρ is the displacement influence factor and is calculated according to equation 3.18: D is the pile diameter; EsL is the Young’s Modulus at the pile tip. [ ] [ ] (3.18) where, νs is the Poisson’s ratio; λ is the lambda factor calculated according to equation 3.19 η1 is the eta factor according to equation 3.20; ξ is the xi factor according to equation 3.21; µL is the mu factor according to equation 3.22; L is the pile length; ρ is the rho factor according to equation 3.23; ζ is the zeta factor according to equation 3.24. ⁄ (3.19) where, Ep is the pile modulus. ⁄ (3.20) where, Db is the diameter of pile base. ⁄ (3.21) Foundation Design of a New Hospital Building in Kecskemét 40 where, Eb is the soil modulus below foundation base. √ (3.22) ⁄ (3.23) where, Esm is the soil modulus at mid-depth of pile shaft. [ [ ] ] (3.24) The Figure 3.8 represents graphically the concept of many of the above mentioned components and also the evolution in depth of the Young’s modulus which is considered to increase linearly along the pile. Figure 3.8 – Young’s modulus vs. depth (Mayne, 2001) Foundation Design of a New Hospital Building in Kecskemét 41 The interaction factor (αij) represents the effect on the global settlement of the pile i due to the load application in another pile (j) within the group (Figure 3.9). Figure 3.9 – Example of interaction factor method (Poulos, 2006) The spacing between the piles (s) emerges, obviously, as the main factor to determine the interaction factor (αij), although the latter is not linearly proportional to the first. These correlations are the aim of the study of several researchers. One of the most recent approaches was proposed by Mandolini and Viggiani (1997). The authors suggested two expressions to define the interaction factor α, the first a polynomial function (3.25) and the second a natural logarithm function (3.26). In order to avoid scale effects, the pile spacing (s) component is normalized by the ratio with the pile diameter (D). ⁄ (3.25) ⁄ (3.26) Where A, B, C and D are verified values tested by Mandolini and Viggiani. Regarding equation 3.25, A ranging between 0.57 to 0.98 whereas B ranges between -0.60 to -1.20. For the second equation (3.26) only one coefficient for each C (1.0) and D (-0.26) was computed. Foundation Design of a New Hospital Building in Kecskemét 42 Foundation Design of a New Hospital Building in Kecskemét 43 4 DESIGN OF THE STUDIED PILED RAFT FOUNDATION 4.1. INTRODUCTION Figure 4.1 represents the detailed structural plan view of the basement of the building “A” which location is defined by the red contour in Figure 2.1.The area of the building is 1059.06 m2 and structural solution is as follows:  64 columns with cross section of 40x40 cm;  4 columns with cross section of 55x55 cm;  6 columns with cross section of 65x65 cm;  Walls of 0.3 m thick. Figure 4.1 – Detail structural plan view (Bohn Deep Foundation Ltd) The location of the piles was defined according to the given geometry (Figure 4.1). Consequently, in the case of a column receiving only one pile, the axis of the latter should overlap the axis of the first one. When the solution is more than one pile per column, the piles must have a distance between axis of three diameters (3D), centered on the column. Foundation Design of a New Hospital Building in Kecskemét 50 Figure 4.5 – Relation between the skin friction and the effective cone resistance (Mahler, 2007) Using the qc,mean values indicated in Table 2.5, Table 2.6 and Table 2.7 and the respective layer thicknesses presented in Table 2.2, Table 2.3 and Table 2.4, it is possible to determine the unit side friction (qsi) as presented in Table 4.4, Table 4.5 and Table 4.6. Table 4.4 – CPT1 qs values for CFA and OMEGA piles qc,mean (MPa) CFA OMEGA αs qs;max (kPa) qsi (kPa) αs qs;max (kPa) qsi (kPa) 26.34 0.55 120 89.3 0.75 160 121.7 3.92 1.00 80 75.2 1.25 100 94.0 9.89 0.55 120 54.7 0.75 160 74.6 1.52 1.00 80 46.7 1.25 100 58.4 10.00 0.55 120 55.0 0.75 160 75.0 Table 4.5 – CPT2 qs values for CFA and OMEGA piles qc,mean (MPa) CFA OMEGA αs qs;max (kPa) qsi (kPa) αs qs;max (kPa) qsi (kPa) 24.44 0.55 120 86.0 0.75 160 117.2 8.35 1.00 80 80.0 1.25 100 100.0 15.39 0.55 120 68.2 0.75 160 93.1 3.16 1.00 80 67.5 1.25 100 84.4 21.64 0.55 120 80.9 0.75 160 110.3 Foundation Design of a New Hospital Building in Kecskemét 51 Table 4.6 – CPT3 qs values for CFA and OMEGA piles qc,mean (MPa) CFA OMEGA αs qs;max (kPa) qsi (kPa) αs qs;max (kPa) qsi (kPa) 7.83 0.55 120 48.7 0.75 160 66.4 1.92 1.00 80 52.6 1.25 100 65.8 13.39 0.55 120 63.6 0.75 160 86.8 The side friction loads (Rs;cal) values are shown in Table 4.7, as result of sum of the product of qsi by the respective shaft area (Asi) in each layer. Table 4.7 – Side friction force (Rs;cal) according to EC7 Rs;cal (kN) CFA OMEGA 0.6 0.8 0.4 0.6 CPT1 1287.62 1716.82 1149.94 1724.90 CPT2 1616.46 2155.28 1424.35 2136.53 CPT3 1390.35 1853.80 1257.95 1886.92 The average, (Rc;cal)mean, and the minimum, (Rc;cal)min, pile compressive resistance values are shown in the Table 4.8. Table 4.8 – Average and minimum values of compressive resistance [kN] CFA OMEGA 0.60 0.80 0.40 0.60 (Rc;cal)mean 2825.62 3843.04 1905.73 3329.83 (Rc;cal)min 2642.61 3649.79 1778.26 3138.62 Finally, in Table 4.9 the design values are calculated, as result of the application of the EC7 suggested method to determine the compressive resistance of a single pile. Table 4.9 – Design total resistance of the piles after the correlation factors [kN] CFA OMEGA 0.60 0.80 0.40 0.60 (Rc;cal)mean 2825.62 3843.04 1905.73 3329.83 ξ3 1.33 1.33 1.33 1.33 (Rc;cal)min 2642.61 3649.79 1778.26 3138.62 ξ4 1.23 1.23 1.23 1.23 Rc;k 2124.52 2889.51 1432.88 2503.63 γt 1.10 1.10 1.10 1.10 Rc;d 1931.39 2626.82 1302.62 2276.03 Foundation Design of a New Hospital Building in Kecskemét 52 4.3.2. DESIGN BEARING CAPACITY ACCORDING TO LCPC METHOD (BUSTAMANTE AND FRANK, 1999) The obtained results for qb for both CFA and OMEGA piles are shown above according to the equation 3.11. The qce values are summarized in Table 4.10. They represent the measured qc between half a diameter (0.5D) above the pile tip and one and a half (1.5D) diameters below. It must be highlighted the high average qc values registered around the base, validating the choice made regarding the pile tip position. Table 4.10 – qce values for the different diameters considered qce (MPa) D=0.4 m D=0.6 m D=0.8 m CPT1 23.302 23.597 23.930 CPT2 25.495 24.706 23.558 CPT3 21.350 21.597 21.284 The value of the Kc takes into account the pile type and the type of the surrounding soil. Bustamante and Frank (1999) suggested, for sand or silty sand, a value of Kc=0.15 for non-displacement (bored or excavated) piles and Kc=0.50 for displacement piles (driven). Since the piles tip is placed in every case in granular soil, the mentioned values were adopted, respectively for the CFA and OMEGA piles. The qb obtained values are presented in Table 4.11. No considerations regarding an upper limit for the pile end resistance were defined by the original authors. Hence, the results are largely influenced by the Kc and by the piles diameter. In fact, the OMEGA piles, in addition to the greater Kc, have a slightly larger average of qce value due to the smaller diameters and consequently more around the base influence. Table 4.11 – Pile end bearing capacity (qb) according to LCPC Method qb (MPa) CFA OMEGA 0.6 0.8 0.4 0.6 CPT1 3.540 3.589 11.651 11.799 CPT2 3.706 3.534 12.748 12.353 CPT3 3.240 3.193 10.675 10.799 The end bearing force values (Rb;cal) are shown in Table 4.12, as result of the multiplication of the qb values by the respective base area (Ab) of the piles. Table 4.12 – Pile end bearing force (Rb;cal) according to LCPC Method Rb;cal (kN) CFA OMEGA 0.6 0.8 0.4 0.6 CPT1 1000.80 1804.24 1464.14 3336.01 CPT2 1047.83 1776.24 1601.90 3492.77 CPT3 915.98 1604.75 1341.48 3053.25 Foundation Design of a New Hospital Building in Kecskemét 53 The obtained results for qs for both CFA and OMEGA piles are shown above according to the equation 3.12. The average CPT cone resistance in each defined layer, so called qc,mean, values are indicated in Table 2.5, Table 2.6 and Table 2.7. The β coefficient is a correlation factor to apply directly to average qc values being defined only for 3 different types of piles (Table 4.13). Table 4.13 – β correlation factor and qs;max based directly on CPT data (Viana da Fonseca et al., 2012) Type of pile Soils Clays and Silts Sands and Gravels Range of qc (MPa) <3 3-6 >6 <5 8-15 >20 Bored without casing (dry method) β - - 75 200 200 200 qs;max (kPa) 15 40 80 40 80 120 Bored with temporary casing β - 100 100 - 100 250 250 300 qs;max (kPa) 15 40 60 40 80 - 40 120 Driven precast concrete β - 75 - 150 150 150 qs;max (kPa) 15 80 80 - - 120 Considering that none of the pile types presented in Table 4.13 match the pile types chosen in the scope of this project, it was necessary the use of the correlation charts (Figure 4.6), converting the CPT qc averages into an equivalent PMT pl. This procedure allows the estimation of qsi based on Pressuremeter test results, wherewith were defined 10 curves relating pl and qs (Figure 4.7) depending on the type of piles (Table 4.17) and the crossed soil (Table 4.18). Figure 4.6 – Equivalent pl depending on the soil type (Bustamante and Gianeselli, 1996) According to Figure 4.6, the ratio qc/pl for clays or clayey silt is 3 and 8 for sand or gravels. In Table 4.14, Table 4.15 and Table 4.16 are presented the equivalent pl associated to each of the identified strata, applying the correspondent ratio for granular and cohesive material. Foundation Design of a New Hospital Building in Kecskemét 54 Table 4.14 – CPT1 equivalent pl values Layer hi (m) hf (m) qc,mean (MPa) pl (MPa) 1 112.40 112.00 26.338 3.29 2 112.00 110.70 3.923 1.31 3 110.70 106.50 9.885 1.24 4 106.50 105.50 1.515 0.51 5 105.50 100.53 9.997 1.25 Table 4.15 – CPT2 equivalent pl values Layer hi (m) hf (m) qc,mean (MPa) pl (MPa) 1 112.40 112.00 24.438 3.05 2 112.00 11.00 8.348 2.78 3 111.00 106.50 15.394 1.92 4 106.50 103.00 3.164 1.05 5 103.00 100.53 21.640 2.71 Table 4.16 – CPT3 equivalent pl values Layer hi (m) hf (m) qc,mean (MPa) pl (MPa) 1 112.40 111.80 7.827 0.98 2 111.80 111.00 1.922 0.64 3 111.00 100.53 13.389 1.67 Foundation Design of a New Hospital Building in Kecskemét 55 As it was referred, the curve Q that connects the pl to qs must be chosen first based on the pile type and after on the soil crossed. Table 4.17 – Description of the 418 analysed piles (Bustamante et al., 2009) Group Code Type No. Piles2 Qty D3 (mm) Depth4 (m) Pile Description 1 1 8 5002000 11.5-23 Pile or Barrette Bored in the Dry 2 64 2701800 6-78 Pile or Barrette Bored with Slurry 3 2 2701200 20-56 Bored and Cased Pile (permanent casing) 4 28 4201100 5.5-29 Bored and Cased Pile (recoverable casing) 51 4 520-880 19-27 Dry Bored Piles or Slurry Bored Piles with Grooved Sockets or Piers (3 types) 2 61 50 410-980 4.5-30 Bored Pile with a single or a double-rotation CFA (2 types) 3 7 48 310-710 5-19.5 Screwed Cast-in-Place 8 1 650 13.5 Screwed Pile with Casing 4 91 30 280-520 6.5-72.5 Pre-cast or Pre-stressed Concrete driven Pile (2 types) 10 15 250-600 8.9-20 Coated Driven Pile (concrete, mortar, grout) 11 19 330-610 4-29.5 Driven Cast-in Place Pile 12 27 170-810 4.5-45 Driven Steel Pile, Closed Tip 5 13 27 190-1.22 8-70 Driven Steel Pile, Open End 6 14 23 260-600 6-64 Driven H Pile 15 4 260-430 9-15.5 Driven Grouted5 or 6 H Pile 7 16 15 - 3.5-2.5 Driven Sheet Pile 1 17 80-140 4-12 Micropile Type I 18 8 120-810 8.5-37 Micropile Type II 8 19 23 1001220 8.5-67 SGP Micropile (Type III) or SGP Pile 20 20 130-660 7-39 MRP Micropile (Type IV) or MRP Pile (1) Some types may include several sub-types. (2) Some piles subjected to several tests. (3) Minimum and maximum nominal diameter B. (4) Minimum and maximum full embedment depth D. (5) Involving a Single Global Post grouting. (6) With Multiple Repeatable Post grouting. Hence, it was considered for the CFA piles the “Bored Pile with a single or a double rotation CFA (2 types)”, corresponding to type 6 and group 2, whereas for the OMEGA piles was selected the “Screwed Cast-in Place Pile” type number 7 and group 3. Based on the above mentioned regarding the characteristic of both piles, the match was clear considering that CFA piles have pile type specifically with its name. The OMEGA piles are part of the “Screw Piles” as defined by the European Nomenclature. Foundation Design of a New Hospital Building in Kecskemét 56 Table 4.18 – Pile type Curves (Qi) associated to a certain soil type (Bustamante et al., 2009) Pile Type No. Clay, Loam Sand, Gravel Chalk Marl, Limestone Weathered Rock 1 Q2 Q2* Q5 Q4 Q6** 2 Q2 Q2 Q5 Q4 Q6** 3 Q1 Q1 Q1 Q2 Q1** 4 Q1 Q2 Q4 Q4 Q4** 5 Q3 Q3* Q5 Q4 Q6** 6 Q2 Q4 Q3 Q5 Q5** 7 Q3 Q5 Q4 Q4 Q4** 8 Q1 Q2 Q2 Q2 Q2** 9 Q3 Q3** Q2 Q2** (a) 10 Q6 Q8 Q7 Q7 (a) 11 Q2 Q3 Q6** Q5** (a) 12 Q2 Q2** Q1 Q2** (a) 13*** Q2 Q1 Q1 Q2 (a) 14*** Q2 Q2 Q1 Q2** (a) 15*** Q6 Q8 Q7 Q7 (a) 16*** Q2 Q2 Q1 Q2** (a) 17 Q1 Q1 Q1 Q2 Q6** 18 Q1 Q1 Q1 Q2 Q6** 19 Q6 Q8 Q7 Q7 Q9** 20 Q6 Q9 Q9 Q9 Q10** * If ground properties permit. ** Use of a higher value must be proven by a load test. *** Cross section and perimeter estimated according to Fig.3. (a) For pile groups No.9 – 16 and if rock condition permits penetration, choose the qs value proposed for marl and limestone or a higher one if this can be proven either by a load test or by reference to an existing example in the same local area The information in Table 4.18 establishes the curve Qi for a certain pile type based on a certain soil crossed. Consequently, “Clay, Loam” corresponds to cohesive layers and “Sand, Gravel” to granular layers. Hence the curves adopted are presented in Table 4.19. Table 4.19 – Respective curves Qi adopted Granular Cohesive CFA Q4 Q2 OMEGA Q5 Q3 Finally the values of qsi are obtained through the analysis of the abacus on Figure 4.7. The results are shown in Table Table 4.20, Table 4.21 and Table 4.22. Foundation Design of a New Hospital Building in Kecskemét 57 Figure 4.7 – Qi Abacus correlating Ménard Limit Pressure (pLM) with skin friction (qs) (Bustamante et al., 2009) Table 4.20 – CPT1 qsi values for CFA and OMEGA piles pl (MPa) CFA OMEGA Tipo qs (kPa) Tipo qs (kPa) 3.29 Q4 170 Q5 200 1.31 Q2 70 Q3 90 1.24 Q4 120 Q5 130 0.51 Q2 30 Q3 50 1.25 Q4 120 Q5 130 Table 4.21 – CPT2 qsi values for CFA and OMEGA piles pl (MPa) CFA OMEGA Tipo qs (kPa) Tipo qs (kPa) 3.05 Q4 160 Q5 190 2.78 Q2 80 Q3 110 1.92 Q4 135 Q5 160 1.05 Q2 55 Q3 75 2.71 Q4 155 Q5 180 Foundation Design of a New Hospital Building in Kecskemét 58 Table 4.22 – CPT3 qsi values for CFA and OMEGA piles pl (MPa) CFA OMEGA Tipo qs (kPa) Tipo qs (kPa) 0.98 Q4 100 Q5 150 0.64 Q2 40 Q3 60 1.67 Q4 130 Q5 180 The last version of the LCPC Method does not provide any guidelines regarding an upper limit for the skin friction (qs;max). However, when correlated with pl, the limiting values of qs are implicitly defined by the evolution of the curves (Figure 4.7). The side friction load (Rs;cal) values are shown in Table 4.23, as result of sum of the product of qsi by the respective shaft area (Asi) of each layer. Table 4.23 – Side friction force (Rs;cal) according to LCPC Method Rs;cal (kN) CFA OMEGA 0.6 0.8 0.4 0.6 CPT1 2430.46 3240.62 1808.43 2712.64 CPT2 2501.05 3334.74 2027.08 3040.62 CPT3 2739.03 3652.04 2503.98 3755.96 The average, (Rc;cal)mean, and the minimum, (Rc;cal)min, pile compressive resistance values are shown in the Table 4.24. Table 4.24 – Average and minimum values of compressive resistance [kN] CFA OMEGA 0.60 0.80 0.40 0.60 (Rc;cal)mean 3545.05 5137.54 3582.33 6463.75 (Rc;cal)min 3431.26 5044.86 3272.56 6048.65 In Table 4.25, the design values are calculated resulting from the application of the LCPC Method. Table 4.25 – LCPC design total resistance of the piles after the correlation factors [kN] CFA OMEGA 0.60 0.80 0.40 0.60 (Rc;cal)mean 3545.05 5137.54 3582.33 6463.75 ξ3 1.33 1.33 1.33 1.33 (Rc;cal)min 3431.26 5044.86 3272.56 6048.65 ξ4 1.23 1.23 1.23 1.23 Rc;k 2665.45 3862.81 2660.62 4859.96 γt 1.10 1.10 1.10 1.10 Rc;d 2423.14 3511.65 2418.74 4418.15 Foundation Design of a New Hospital Building in Kecskemét 59 The Table 4.26 summarizes the design values for the ultimate pile bearing capacity, where the LCPC proved to be less conservative than the EC7 Approach. Hereinafter, the design, including the determination of the number of piles, shall be made based on the latter. Table 4.26 – Comparison between the Rc;d obtained by the two methods Rc;d (kN) CFA OMEGA 0.6 0.8 0.4 0.6 EC7 Approach 1931.39 2626.82 1302.62 2276.03 LCPC Method 2423.14 3511.65 2418.74 4418.15 4.3.3. DTU 13.2 NORM (1992) The French Norm DTU 13.2 “Foundations Profondes pour le Bâtiment” established a criterion for the pile design, apart from the outlined methods that estimate “compressive resistance of the ground against a pile, at the ultimate limit state” (in Eurocode 7). For reinforce concrete piles, the criteria controls the stress level of concrete avoiding the steel yielding and concrete cracking. The approach defines a conventional concrete compressive strength, fc* (equation 4.2): (4.2) where, fcj is the characteristic compressive cylinder strength of concrete at j days, according to BAEL guidelines; fclim is a limit strength of concrete depending on pile foundation type (Table 4.27); k1 is a coefficient that takes into account the execution method of the pile (Table 4.27); k2 is a coefficient that takes into account the difficulty related to the concrete casting of the pile. Table 4.27 – Relation between the pile type and the respective fclim and k1 (DTU 13.2, 1992) Group Pile Foundation Type fclim k1 A Borehole piles or precast walls fc28 1.00 Pre-stressed tubular piles fcj 1.15 Prefabricated piles dynamically driven fcj 1.15 Well with vibrated concrete fc28 1.00 Well with self-compacting concrete fc28 1.20 B Driven Cast-in Place Pile fc28 1.30 Pile or Barrette Bored in the dry fc28 1.30 Bored and Cased Pile  Dry casting fc28 1.20  Casting underwater fc28 1.30 Pile and Barrette Bored with Slurry (generally bentonite) fc28 1.30 Foundation Design of a New Hospital Building in Kecskemét 66 Foundation Design of a New Hospital Building in Kecskemét 67 5 ASSESSMENT OF THE SETTLEMENTS 5.1. INTRODUCTION After reaching the optimal pile foundation system, the settlement assessment was carried out only considering the OMEGA piles and the geometry defined previously (Subchapter 4.4). The deformation analysis was made following two methods of different nature:  Soil-structure modelling using AxisVM  Interaction Factor Method (Poulos and Davis, 1980). All the theoretical background that led to the application of the referred methods was presented in detail in Subchapter 3.3. Therefore, this chapter is firstly intended to present the obtained results followed by the discussion and comparison of them. 5.2. SOIL-STRUCTURE MODELLING USING AXISVM The program AxisVM is a structural analysis and design software (http://axisvm.eu) being developed by InterCAD Kft. The company is based Budapest, Hungary since 1991. The software used in the scope of this work, AxisVM 11®, is a three-dimensional program based on the finite element method (FEM). Being a purely structural program, the modelling of this specific geotechnical problem was done through the definition of stiffness parameters that effectively portray the deformability behavior of the piles and the soil. Thus, the assignment of spring constants permitted the simulation of piles and the definition of a continuous 2D element represented the soil stiffness. 5.2.1. PILE FLEXIBILITY ESTIMATION The spring constant can be defined in the software as a nodal support (Figure 5.1) considering that the pile is a one-dimensional element (FL-1). This value corresponds to the pile flexibility and was determined through the theoretical load-settlement curves, as follows. Foundation Design of a New Hospital Building in Kecskemét 68 Figure 5.1 – “Nodal support” option in “Elements” menu (adapted from AxisVM) Assuming the shaft resistance mobilization for 1.5% of pile diameter and the base resistance mobilization for 10% of pile diameter (Table 5.1), the load-settlement curve was defined, for both diameter of 0.4 m and 0.6 m, as the sum of the two components. Table 5.1 – Settlements associated to resistance mobilization s (m) D (m) 0.40 0.60 0.015D 0.006 0.009 0.1D 0.040 0.060 Retrieving the Table 4.3 and Table 4.7 with the end bearing force and the side friction force data, respectively and linking the OMEGA values with the settlements established in Table 5.1, it is possible to define a load-settlement curve for each of the CPTU tests studied. Table 5.2 and Table 5.3 present the values of the load-settlement curves for the diameter of 0.4 m and 0.6 m, respectively. Table 5.2 – Load-settlement curve data for 0.4 m diameter pile s (m) Rc (kN) CPT1 CPT2 CPT3 Average 0.000 0.00 0.00 0.00 0.00 0.006 1244.18 1518.60 1352.19 1371.66 0.040 1778.26 2052.67 1886.27 1905.73 Table 5.3 – Load-settlement curve data for 0.6 m diameter pile s (m) Rc (kN) CPT1 CPT2 CPT3 Average 0.000 0.00 0.00 0.00 0.00 0.009 1936.96 2348.58 2098.98 2128.17 0.060 3138.62 3550.24 3300.64 3329.83 Associated to the previous tables, the Figure 5.2 and Figure 5.3 illustrate the load-settlement curves for each of the CPTU tests and the average of those results. Foundation Design of a New Hospital Building in Kecskemét 69 Figure 5.2– Load-settlement curve for 0.4 m diameter pile Figure 5.3 – Load-settlement curve for 0.6 m diameter pile Considering the forces applied to the columns base and subsequently the forces applied to piles bases, presented in Appendix A, the maximum characteristic loads applied on the piles are summarized in Table 5.4: Table 5.4 – Maximum characteristic loads D (m) 0.4 0.6 Fk,max (kN) 817.87 1604.85 Therefore it can be concluded that all the loads applied on the piles belong to the first line of the theoretical load-settlement curve reflecting that the pile flexibility was simply determined according to equation 5.1: 0.000 0.010 0.020 0.030 0.040 0.050 0.060 0.070 0.080 0.090 0.00 1000.00 2000.00 s (m) R (kN) CPTU1 CPTU2 CPTU3 Average 0.000 0.010 0.020 0.030 0.040 0.050 0.060 0.070 0.080 0.090 0.00 1000.00 2000.00 3000.00 4000.00 s (m) R (kN) CPTU1 CPTU2 CPTU3 Average Foundation Design of a New Hospital Building in Kecskemét 70 (5.1) The input parameters introduced in AxisVM are presented in Table 5.5. Table 5.5 – Pile flexibility for each diameter D (m) 0.4 0.6 kP (MN/m) 228.61 236.46 5.2.2. DETERMINATION OF MODULUS OF SUBGRADE REACTION The continuous 2D surface support (Figure 5.4) introduced in the program is in fact the modulus of subgrade reaction (Lopes, 2001) or the Winkler (1867) spring stiffness being determined by the equation 5.2: Figure 5.4 – “Surface support” option in “Elements” menu (adapted from AxisVM) (5.2) where, p is the raft distributed load; w is the raft settlement. The distributed load, p, corresponds exactly to the sum of all the characteristic values divided by the raft area and is equal to 107.06 kN/m2. The raft settlement was determined through the Schmertmann method considering an equivalent rectangular raft area (Table 5.6) and the Young’s Modulus evolution with depth as defined in equation 2.14. Since the method was developed to predict settlements on sands (which represent 80% of the studied soil profile), the consolidation effects associated to the clayey layers were neglected. However, this assumption does not affect significantly the obtained results. Table 5.6 – Equivalent rectangular raft dimensions L (m) B (m) L/B A (m2) 52 20.35 2.56 1058.2 Foundation Design of a New Hospital Building in Kecskemét 71 The calculation of the vertical strain influence factor, Iε, was done following the Table 3.4 for intermediate L/B between 1 and 10, in which it is proposed an equation that establishes the interpolation. Hence, for each sublayer considered the value Iε for axisymmetric and for strip (continues) foundation taking the respective Iεp. For the latter, it was considered the unit weights defined in 2.6.1 and the adopted stratigraphy presented in Table 2.2, Table 2.3 and Table 2.4 in order to calculate the effective vertical stress. The results are presented in the appendix (A.2) to this thesis. The vertical strain influence factor evolution with depth is presented in Figure 5.5. Figure 5.5 – Iε depth evolution for L/B=2.56 Analysing Figure 5.5, it can be observed that the specific case study (L/B=2.56) is characterized by a slightly smaller peak of Iε but, on other hand, by a larger starting Iε (0.12) and a larger influence depth (z/B=2.18) when compared with axisymmetric foundation. The obtained raft settlements are resumed in Table 5.7, with reduced variations between each of the tests. Table 5.7 – Raft settlement CPTU s (mm) 1 12.45 2 12.45 3 12.36 Consequently, the values for the modulus of subgrade reaction are also close to each other (Table 5.8). For the AxisVM modelling the average value should be computed. 0.00 0.50 1.00 1.50 2.00 2.50 3.00 3.50 4.00 4.50 0.000 0.200 0.400 0.600 0.800 z/B Iε Axisymmetric Continues 1<L/B<10 Foundation Design of a New Hospital Building in Kecskemét 72 Table 5.8 – Modulus of subgrade reaction CPTU k (kN/m3) 1 8602.57 2 8599.39 3 8662.14 Average 8621.36 Taking into account the simplified calculation carried out for both the pile flexibility and the modulus of subgrade reaction, it was considered appropriate to vary slightly the average values of the referred stiffness parameters. Hence, the following parametric analyses were the defined, in addition to the use of the average values:  20% decrease of the pile stiffness (0.8.kp) and 20% increase of the modulus of subgrade reaction (1.2.ks);  20% increase of the pile stiffness (1.2.kp) and 20% decrease of the modulus of subgrade reaction (0.8.ks);  20% decrease of the pile stiffness (0.8.kp) and 20% decrease of the modulus of subgrade reaction (0.8.ks); The objective of theses analyses is to study the input parameters influence in the maximum settlements generated, maximum differential settlement and angular distortion. The Table 5.9 summarizes the nodal and the surface supports associated to each of the 4 studies carried out (including the average and the variations from this). Table 5.9 – Stiffness parameters input Case Study kp (kN/m) ks (MN/m3) D=0.4 D=0.6 1 228609.84 236463.82 8621.36 2 182887.87 189171.05 10345.64 3 274331.80 283756.58 6897.09 4 182887.87 189171.05 6897.09 5.2.3. MODEL CONSTITUTION As an initial iteration, a raft of 0.8 m thick of C25/30 concrete class was considered. However, due to the significant discrepancy between the applied loads in different raft zones, an efficient solution to create greater concrete thickness (1.0 m) in the most loaded areas was considered. This decision is conservative in view of the subsequent reinforcement design, reducing the steel ratio. In order to design a more rigid structure, in addition to the inclusion of 74 columns, it was also modelled the first floor slab of 20 cm thick. The Figure 5.6 exemplifies the considered model. Foundation Design of a New Hospital Building in Kecskemét 73 Figure 5.6 – 3D perspective of the defined model (adapted from AxisVM) In the 3D wireframe (Figure 5.6), the columns are countered by blue and the piles are represented in brown. For the settlement assessment, the reinforcement is not relevant so it was not specified for this calculation. The mesh is defined by elements of 0.5 m size which can reproduce, considering the significant dimensions of the structure to model, its deformation behaviour fundamentally correct. The model run was performed through a linear static analysis. 5.2.4. AXISVM RESULTS Taking into account the dimensions of the modelled structure, it becomes impossible to present and study all results obtained by the program. Thus, it is essential to select the most significant results in the scope of this project. When assessing the piled raft settlements, results such as the pile settlement and the differential settlements between piles are very important, as stated above. Figure 5.7, Figure 5.8, Figure 5.9 and Figure 5.10 represent the 2D isosurfaces obtained by the program associated to displacement in vertical direction (eZ) for each of the 4 cases studied. Figure 5.7 – Case study 1 Isosurface 2D (adapted from AxisVM) Foundation Design of a New Hospital Building in Kecskemét 74 Figure 5.8 – Case study 2 Isosurface 2D (adapted from AxisVM) Figure 5.9 – Case study 3 Isosurface 2D (adapted from AxisVM) Figure 5.10 – Case study 4 Isosurface 2D (adapted from AxisVM) As can be seen, the simple observation of the figures above does not allow inferring relevant conclusions regarding the variations imposed. In fact, only slightly differences in the isosurfaces boundaries can be noticed. Thus, it was necessary, in order to understand the relative influence of this Foundation Design of a New Hospital Building in Kecskemét 75 binomial, to analyse specifically the obtained values. Thus, there were interpreted the results for the settlements (s) of each pile and the differential settlement (Δs) and the angular distortion (α= Δs/L) between every pair of piles (separated by L meters), summarized in Table 5.10, Table 5.11 and Table 5.12. Table 5.10 – Maximum and average settlements Case Study 1 2 3 4 smax (mm) 4.843 5.649 4.230 6.000 savg (mm) 3.555 3.957 3.236 4.425 smax/savg 1.362 1.428 1.307 1.356 Table 5.11 – Maximum and average differential settlements Case Study 1 2 3 4 Δsmax (mm) 1.912 2.521 1.575 2.365 Δsavg (mm) 0.568 0.740 0.474 0.706 Δsmax/Δsavg 3.366 3.407 3.323 3.350 Table 5.12 – Maximum and average angular distortions Case Study 1 2 3 4 αmax 1/3597 1/2421 1/4115 1/3049 αavg 1/29412 1/22222 1/34483 2/23256 Through the observation of Table 5.10, it can be concluded that the settlement magnitude is mostly defined by the pile flexibility, when comparing the fluctuations establishing the case study 4 as the base. When increasing the soil spring constant from 0.8ks to 1.2ks (case study 2) while maintaining the pile flexibility, the maximum settlement (produced in any case in the same pile, over the column D5) is just reduced in 6%, whereas the increase of the pile flexibility from 0.8ks to 1.2ks while maintaining the soil spring constant induces a settlement reduction of 30%. Logically, the higher the ratio smax/savg the greater the differential settlement, fact corroborated in Table 5.11. Other important aspect resulting of the comparison between the case study 2 and 3 is that the variation of the greater magnitude parameter will control the differential settlement ratio between the maximum and the average. Regarding the angular distortions, the case study 2 resulted in the largest percentages but still well below the regulatory limits. Foundation Design of a New Hospital Building in Kecskemét 82 Table 6.1 – Minimum pile reinforcement area D (m) As,bpmin 0.4 7Φ16 0.6 10Φ16 Regarding the transverse reinforcement, EN 1536 suggests, for an helical reinforcement: { } (6.2) Thus, the diameter for the transverse reinforcement should be 6 mm and the spacing shall not be less than the clear distance obtained in the longitudinal reinforcement (180 mm). 6.3. RAFT FOUNDATION DESIGN Being a structural software, highly targeted to solve project queries, AxisVM enables a failure calculation, which leads to the reinforced concrete design. The model used to run this calculation is the same used to the settlements assessment, illustrated in Figure 5.6 as well as the geometry of the structural elements defined in subchapter 5.2.3. The reinforced slab to design is represented in Figure 6.1 as 3D perspective with the 0.8 m raft and the 6 other zones with 1.0 m, identifiable by its salience. The reason for the extra concrete thickness zones is related to the significant applied loads over these areas, as it was pointed out in the previous chapter. Figure 6.1 – 3D perspective of the raft foundation (adapted from AxisVM) The concrete class of the slab is C25/30, mentioned above in subchapter 5.2.3, the rebar steel is S500 and associated with the adopted geometry complete the material input. The model adopted was a linear static type analysis but considering a different load case from the previous analysis. This time, the ultimate limit state was used considering the design values of the applied loads over the piles. The software AxisVM also enables the option of performing the reinforcement calculations under the Eurocode 2 regulations, which was then the code adopted. The “R.C. Design” menu on AxisVM comprises the “Reinforcement Parameters” and the “Actual Reinforcement”. The first is set to introduce the thickness of the material (slab) and its mechanical parameters, in the case of being reinforced concrete, both the concrete and steel classes are defined. Foundation Design of a New Hospital Building in Kecskemét 83 The second is set to define a certain level of reinforcement for the modelled elements, specifically the slab. Focusing on the actual reinforcement, it should be noted that the software requires the definition of the rebars cover on top and bottom surfaces, which was set as the minimum (26 mm) as shown in Figure 6.2. Figure 6.2 – Cover to reinforcement (adapted from AxisVM) The actual reinforcement is designed in directions x and y, according to Figure 6.3, and on top and bottom of the cross section (Figure 6.4). Figure 6.3 – Plan xOy (adapted from AxisVM) Foundation Design of a New Hospital Building in Kecskemét 84 Figure 6.4 – Actual reinforcement window on AxisVM software (adapted from AxisVM) The level of required reinforcement is defined comparatively with the “reinforcement values” provided by the program, ensuring that the difference between the two is positive. Obviously, the smaller the reinforcement differences the more economic is the solution. To perform this calculation, it was chosen the Case Study 1 (see 5.2.2) that represents the average values for the pile flexibility (kp) and the soil stiffness (ks). The results obtained via AxisVM are summarized in Table 6.2, where the reinforcement is indicated according to the direction and the position. Distinction between the 6 thicker zones (numbered clockwise) and the main slab is made too. Table 6.2 – Required reinforcement for the raft design Direction Position Extra thick zones Main slab 1 2 3 4 5 6 Φ (mm) s (mm) Φ (mm) s (mm) Φ (mm) s (mm) Φ (mm) s (mm) Φ (mm) s (mm) Φ (mm) s (mm) Φ (mm) s (mm) x t 16 330 16 360 16 360 16 420 16 420 16 350 16 100 b 25 180 25 170 25 170 25 160 25 160 25 170 25 200 y t 16 260 16 240 16 230 16 330 16 380 16 360 16 230 b 25 180 25 180 25 180 25 160 25 160 25 160 25 230 *t: top, b: bottom Observing the Table 6.2, it is possible to conclude that the creation of the extra thick zones led to the decrease of the steel reinforcement ratio in both the top and the bottom surfaces, considering that even being 20 cm thicker, the reinforcement level is still higher than in the main slab. Moreover, the increase of the slab thickness in certain critical zones permitted a more leveled amount of reinforced concrete all over the raft. To complement the Table 6.2, it was considered suggestive the presentation of the reinforcement differences 2D isosurfaces of the top surface x direction (xt), the bottom surface x direction (xb), the Foundation Design of a New Hospital Building in Kecskemét 85 top surface y direction (yt), the bottom surface y direction (yb), (Figure 6.5, Figure 6.6, Figure 6.7 and Figure 6.8). Figure 6.5 – Reinforcement difference xt (adapted from AxisVM) Figure 6.6 – Reinforcement difference xb (adapted from AxisVM) Figure 6.7 – Reinforcement difference yt (adapted from AxisVM) Foundation Design of a New Hospital Building in Kecskemét 86 Figure 6.8 – Reinforcement difference yb (adapted from AxisVM) Figure 6.5, Figure 6.6, Figure 6.7 and Figure 6.8 show the reinforcement differential in both x and y direction combined with the top and the bottom surfaces. According to the colour legend on the right, the more red the scale is the larger the reinforcement differential. The reinforcement design was carried out assuring that no node, within any of the 7 domains, would have a negative difference between the required and the actual reinforcement. Hence, this permitted the optimization of the more confined domains (6 thicker slabs) and led, at the same time, to a noneconomic design of the larger domain (main slab). In order to optimize the latter, more discretized domains should have been created, approaching the differences to zero and eliminating the red coloured surfaces. Foundation Design of a New Hospital Building in Kecskemét 87 7 DESIGN OF THE EXCAVATION DEWATERING SYSTEM 7.1. INTRODUCTION The groundwater on a particular construction site is caused essentially by heavy rainfall and percolation through permeable soils. The permeability of the soils will govern the water table (or phreatic level) position in the subsoil and its variations. Within a Civil Engineering work, the reduction of the groundwater level is crucial considering the effects the water has on the strength properties of the soil and the associated deformations that can be induced by the pore water pressure variation, which is the same as the effective stress changes. However in the specific case of a shallow excavation, the increase of soil strength, induced by the reduction of the pore water pressure, cannot be considered relevant when compared with its effects in a deep excavation execution. In the scope of a shallow excavation execution (Figure 7.1 – Water conditions in the construction site), the high position of the water table in situ is more directly affected by the construction methods and consequently its duration and cost. In a long term analysis, the durability of the built structure could be affected and the maintenance costs can be significant (Puller, 2003). Therefore, the control of groundwater gains particular relevance in the geotechnical project planning. Figure 7.1 – Water conditions in the construction site In relation to Figure 7.1, the excavation was considered to be 4.50 m, slightly above the measured values in the investigation drillings (see 2.1), fact that will not influence the results, though. Foundation Design of a New Hospital Building in Kecskemét 88 This chapter addresses a possible temporary dewatering system for the executed excavation (Figure 7.1) by creating wells using the program Plaxis 2D and a classical empirical formulation proposed by Mansur and Kaufman (1962). To run the calculations, it was adopted a simplified stratigraphy based on the 3 CPTU tests as shown in Figure 7.2. Figure 7.2 – Adopted stratigraphy for the groundwater calculations 7.2. ASSESSMENT OF THE COEFFICIENT OF PERMEABILITY Recalling the adopted soil stratification, two main soil behaviour types were defined, granular and cohesive. Hence, it was necessary to arbitrate the value of the soil coefficient of permeability (k), considering common sands and clays. In the absence of the grain size distribution curve, which is remarkably the most relevant element for the evaluation of this parameter, the permeability (also called hydraulic conductivity) estimative was done using standard values available in the literature (Matos Fernandes, 2006). For the granular soils, it was considered k=10-5 m/s, whereas for the cohesive strata, it was considered k=10-8 m/s. 7.3. PLAXIS 2D MODELLING The program Plaxis 2D is a Dutch software based on the finite element method intended to solve various geotechnical problems, including hydraulic coupling. The modelling in the program was done according to Figure 7.2, assigning, according to the strata, its respective characteristics, which, in this case, are the hydraulic conductivity in direction x (kx) and y (ky) since it is related to a percolation problem. In addition to this, two narrow trenches (0.50 m each) were created with no domain assigned to represent the wells. The option for this procedure instead of using the Well feature offered by the program was due to the boundary condition problems related to cross different layers. The Figure 7.3 represents the modelled structure in Plaxis 2D. Foundation Design of a New Hospital Building in Kecskemét 89 Figure 7.3 – Modelled structure in Plaxis 2D (adapted from Plaxis 2D) In order to accurately portray the groundwater conditions far from the excavation pit, a wide mesh was created, assuring that near the lateral boundaries the water table remains in the initial level. Moreover, this covers all the variations from the initial conditions until the final lowered water table, close to the wells. The mesh is composed by triangular elements (Figure 7.4), defined as “Coarse” in the “Global Coarseness” menu. Figure 7.4 – Generated mesh in Plaxis 2D (adapted from Plaxis 2D) The definition of the water conditions was done through the assignment of the boundary type “Head (user-defined)”, where it was necessary to insert, for each of the lateral boundary points, the piezometric head, hw, in relation to the phreatic level. The Figure 7.5 illustrates the “Boundary Conditions” window. Foundation Design of a New Hospital Building in Kecskemét 90 Figure 7.5 – Boundary conditions window in Plaxis 2D (adapted from Plaxis 2D) The drainage system was simulated using the “Flow Mode” and the calculation type “Groundwater flow – steady state”. The length of the trenches was iterated in order to achieve at least a difference of 1 meter between the groundwater table and the bottom of the excavation (working level), a distance considered reasonable to guarantee satisfactory working conditions and the economy of the well execution. The optimal solution was found with a pair of trenches with 7.5 m depth and 3.0 m below the working level as shown Figure 7.6. The coloured shades represent the different groundwater heads in the different nodes of the modelled mesh. Figure 7.6 – Evolution of the water table position (adapted from Plaxis 2D) This configuration of the trenches places the water table 1.30 m (z=17.20 m) below the working level (Figure 7.7) from an initial position at 2.95 m, which means a lowering of 4.25 m. Foundation Design of a New Hospital Building in Kecskemét 91 Figure 7.7 – Lowering level obtained with 7.5 length trenches (adapted from Plaxis 2D) Between the trenches and the slopes, arose some irregular lines, displaying the same label as the water table that the author believes to be due to numeric errors. Regarding the flow values around the bottom of the trenches, through the stress points it was possible to obtain the absolute mode flow as show in Figure 7.8 and Figure 7.9 and summarized in Table 7.1. Figure 7.8 – Left trench |q| values (m3/day/m) (adapted from Plaxis 2D) Figure 7.9 – Right trench |q| values (m3/day/m) (adapted from Plaxis 2D) Foundation Design of a New Hospital Building in Kecskemét 98 Junttan. Respecting Ground (2013). http://www.junttan.com/piling_solutions (Accessed in March 2013). Kulhawy, F.H. and Mayne, P.W. (1990). Manual on Estimating Soil Properties for Foundation Design. Geotechnical Engineering Group, Cornell University. Ithaca, New York, USA. Lakatos, I. (2011). Cölöpalapok méretezése az Eurocode 7 követelményei szerint. PhD Thesis, BME. Budapest, Hungary. (in Hungarian) Lopes, F.R. (2000). Design of raft foundations on Winkler springs. Vol. Design applications of raft foundations. Lunne, T., Robertson, P.K. and Powell, J.J.M. (1997). Cone Penetration Testing in Geotechnical Practice. Blackie Academic & Professional. London, United Kingdom. Mahler, A. (2007). Settlement Prediction of CFA Piles Based on CPTu Results. BME. Budapest, Hungary. Mandolini, A. and Viggiani, C. (1997). Settlement of Piled Foundations. Geotechnique no. 47 (4), 791-816. Mansur, C. and Kaufman, R. (1962). Dewatering. Foundation engineering. Foundation Engineering, 241-350. Matos Fernandes, M. (2006). Mecânica dos Solos – Conceitos e Princípios Fundamentais. FEUP edições. Porto, Portugal. (in Portuguese) Matos Fernandes, M. (2011). Mecânica dos Solos – Introdução à Engenharia Geotécnica. FEUP edições. Porto, Portugal. (in Portuguese) Mayne, P.W. (2007). Cone Penetration Testing State-Of-Practice. Transportation Research Board, Vol. 368. Mayne, P.W. and Schneider, J.A. (2001). Evaluating Axial Drilled Shaft Response by Seismic Cone. Foundations & Ground Improvement, GSP 113, ASCE, 655-669. Poulos, H.G. (1987). From theory to practice in pile design. Vol. 55, No. 9. Poulos, H.G. (2006). Pile Group Settlement Estimation – Research To Practice. Geotechnical Special Publication 153. Poulos, H.G. and Davis, E.H. (1980). Pile Foundation Analysis and Design. No. Monograph. Puller, M. (2003). Deep Excavations: a practical manual. Thomas Telford, ASCE. Robertson, P.K. (1990). Soil classification using the cone penetration test. Canadian Geotechnical Journal, 27(1), 151-158. Robertson, P.K. (2009). Interpretation of cone penetration tests-a unified approach. Canadian Geotechnical Journal 46.11, 1337-1355. Robertson, P.K. (2012). Interpretation of in-situ tests – some insights. Mitchell Lecture - ISC'4 Brazil, Sept., 2012. Gregg Drilling & Testing, Inc. Signal Hill, California, USA. Robertson, P.K., and Campanella, R.G. (1983). Interpretation of cone penetration tests – Part I: sand. Canadian Geotechnical Journal, 20(4), 718-733. Robertson, P.K. and Cabal, K.L. (2012). Guide to Cone Penetration Testing for Geotechnical Engineering. Gregg Drilling & Testing, Inc. Signal Hill, California, USA. Foundation Design of a New Hospital Building in Kecskemét 99 Robertson, P.K., Campanella, R.G., Gillespie, D. and Greig, J. (1986). Use of piezometer cone data. Proceedings of the ASCE Speciality Conference In Situ ’86: Use of In Situ Tests in Geotechnical Engineering, ASCE, 1263-1280. Blacksburg, Virginia, USA. Rocscience (2013). http://www.rocscience.com/settlecalc/schmerthelp.html (Accessed in April 2013). Schmertmann, J.H. (1970). Static cone to compute settlement over sand. Journal of the Soil Mechanics Engineering Division, ASCE, Vol. 96, No. 3, 1011-1043. Senneset, K. and Janbu, N. (1985). Shear strength parameters obtained from static cone penetration tests. Strength Testing of Marine Sediments: Laboratory and In-Situ Measurements. Symposium, ASTM Special technical publication, STP 883, 41-54. Senneset, K. and Nestvold, J. (1992). Deep Compaction by Vibro Wing Technique and Dynamic Compaction. Geotechnical Special Publication, ASCE, Vol. 2, No. 30, 889-901. Viana da Fonseca, A. and Santos, J. (2008). International Prediction Event on the Behaviour of Bored, CFA and Driven Piles in CEFEUP/ISC’2 Experimental Site – 2003. Final Report. FEUP edições. Porto, Portugal. Viana da Fonseca, A, Buttling, S. and Coutinho, R.Q. (2012). Foundations: Shallow and Deep Foundations, Unsaturated Conditions, Heave and Collapse, Monitoring and Proof Testing. Vol. Handbook of Tropical Residual Soils Engineering. CRC Press. London, United Kingdom. Winkler, E. (1867). Die Lehre von der Elastizitat und Festigkeit. Dominicus. Prague, Czech Republic. (in German) Zhang, G., Robertson, P.K. and Brachman, R.W.I. (2002). Estimating liquefaction-induced ground settlements from CPT for level ground. Canadian Geotechnical Journal, 39(5), 1168–1180. Foundation Design of a New Hospital Building in Kecskemét 100 Foundation Design of a New Hospital Building in Kecskemét 101 APPENDIX A A.1. CHARACTERISTIC AND DESIGN LOADS Table A.1 – Piles data Column Code Pile Number Pile Diameter x (m) y (m) Ed (kN) Ek (kN) A5 1 0,60 12,23 20,90 1616,31 1154,51 A6 1 0,60 15,84 20,90 1832,29 1308,78 A7 1 0,60 19,44 20,90 1718,32 1227,37 A8 1 0,60 23,04 20,90 1922,30 1373,07 A9 1 0,60 26,63 20,90 1738,24 1241,60 A10 1 0,60 30,23 20,90 2066,09 1475,78 A11 1 0,60 33,83 20,90 1922,43 1373,16 A12 1 0,60 37,43 20,90 2157,01 1540,72 A13 1 0,60 41,03 20,90 1916,63 1369,02 A14 1 0,60 44,63 20,90 1976,63 1411,88 A15 1 0,60 48,23 20,90 1588,82 1134,87 B1 1 0,60 0,20 20,34 1338,44 956,03 B2 1 0,60 3,50 20,34 1626,28 1161,63 B4 1 0,60 8,15 20,34 1851,85 1322,75 B5 1 0,60 12,23 20,34 1740,94 1243,53 B16 1 0,60 50,03 20,34 1753,83 1252,74 B17 1 0,60 51,83 20,34 1437,91 1027,08 C11 1 0,60 0,20 17,71 1339,41 956,72 C12 1 0,60 3,50 17,71 1337,70 955,50 C13 1 0,60 6,30 17,34 1276,78 911,99 C14 1 0,60 7,67 17,34 1520,65 1086,18 C15 1 0,60 12,23 17,19 1952,84 1394,89 C21 1 0,60 0,20 14,25 1377,66 984,04 C22 1 0,60 3,50 14,25 1454,28 1038,77 C23 1 0,60 6,30 14,25 1543,56 1102,54 C24 1 0,60 7,67 14,25 1493,77 1066,98 C25 1 0,60 12,23 14,25 2246,79 1604,85 C26 1 0,60 14,94 14,20 1860,26 1328,75 2 0,60 16,74 14,20 1860,26 1328,75 Foundation Design of a New Hospital Building in Kecskemét 102 C28 1 0,60 22,14 15,10 1424,60 1017,57 2 0,60 23,94 15,10 1424,60 1017,57 3 0,60 23,94 13,30 1424,60 1017,57 4 0,60 22,14 13,30 1424,60 1017,57 C210 1 0,60 29,33 15,10 1425,60 1018,28 2 0,60 31,13 15,10 1425,60 1018,28 3 0,60 31,13 13,30 1425,60 1018,28 4 0,60 29,33 13,30 1425,60 1018,28 C212 1 0,60 36,53 15,10 1421,96 1015,69 2 0,60 38,33 15,10 1421,96 1015,69 3 0,60 38,33 13,30 1421,96 1015,69 4 0,60 36,53 13,30 1421,96 1015,69 C214 1 0,60 43,73 14,20 1867,35 1333,82 2 0,60 45,53 14,20 1867,35 1333,82 C215 1 0,60 47,33 14,24 1161,20 829,43 2 0,60 49,13 14,24 1161,20 829,43 C217 1 0,60 51,83 14,24 1473,25 1052,32 C31 1 0,60 0,20 8,34 1549,27 1106,62 C32* 1 0,60 4,45 8,34 1506,02 1075,73 C33 1 0,60 6,30 8,34 1763,11 1259,36 C34 1 0,60 7,67 8,34 1727,86 1234,19 C35 1 0,60 12,23 8,34 2180,58 1557,56 C41 1 0,60 0,20 5,27 1517,45 1083,89 C42* 1 0,60 4,45 5,27 1507,82 1077,01 C43 1 0,60 6,30 5,27 1379,09 985,06 C44 1 0,60 7,67 5,27 1464,25 1045,89 C45 1 0,60 12,23 5,27 1939,73 1385,52 C51 1 0,60 0,20 2,24 1558,82 1113,44 C52* 1 0,60 4,45 2,24 1549,35 1106,68 C54 1 0,60 7,67 2,24 1626,74 1161,96 C55 1 0,60 12,23 2,24 1927,27 1376,62 D5 1 0,60 12,23 0,20 1768,92 1263,51 D6 1 0,60 15,84 0,20 1980,25 1414,46 D7 1 0,60 19,44 0,20 2058,84 1470,60 D8 1 0,40 22,44 0,20 1145,02 817,87 2 0,40 23,64 0,20 1145,02 817,87 D9 1 0,60 26,63 0,20 2099,97 1499,98 D10 1 0,60 29,33 0,20 1215,21 868,01 2 0,60 31,13 0,20 1215,21 868,01 D11 1 0,60 33,83 0,20 2078,18 1484,41 D12 1 0,60 37,43 0,20 2195,62 1568,30 D13 1 0,60 41,03 0,20 1886,86 1347,76 D14 1 0,60 44,63 0,20 1766,67 1261,91 D15 1 0,60 48,23 0,20 1324,92 946,37 E117 1 0,60 51,83 17,19 1453,01 1037,86 E415 1 0,60 48,23 9,31 1423,57 1016,84 E416 1 0,60 50,03 9,31 1289,41 921,01 E417 1 0,40 51,83 9,31 1051,28 750,91 E56 1 0,60 14,94 7,80 1922,53 1373,24 2 0,60 16,74 7,80 1922,53 1373,24 Foundation Design of a New Hospital Building in Kecskemét 103 E58 1 0,60 22,14 8,70 1486,18 1061,55 2 0,60 23,94 8,70 1486,18 1061,55 3 0,60 23,94 6,90 1486,18 1061,55 4 0,60 22,14 6,90 1486,18 1061,55 E510 1 0,60 29,33 8,70 1503,17 1073,69 2 0,60 31,13 8,70 1503,17 1073,69 3 0,60 31,13 6,90 1503,17 1073,69 4 0,60 29,33 6,90 1503,17 1073,69 E512 1 0,60 36,53 8,70 1506,98 1076,41 2 0,60 38,33 8,70 1506,98 1076,41 3 0,60 38,33 6,90 1506,98 1076,41 4 0,60 36,53 6,90 1506,98 1076,41 E514 1 0,60 43,73 7,80 1927,94 1377,10 2 0,60 45,53 7,80 1927,94 1377,10 E615 1 0,60 48,23 6,36 1713,04 1223,60 E616 1 0,60 50,03 6,36 1290,34 921,67 E617 1 0,60 51,83 6,36 1326,40 947,43 E715 1 0,60 48,23 2,00 1417,32 1012,37 E716 1 0,60 50,03 2,00 1435,18 1025,13 E717 1 0,60 51,83 2,00 1333,55 952,54 Foundation Design of a New Hospital Building in Kecskemét 104 Figure A.1 – Representation of the design loads over the respective column Foundation Design of a New Hospital Building in Kecskemét 105 A.2. SCHMERTMANN METHOD RESULTS Table A.2 – Schmertmann Method results for CPT1 z (m) z/B γ (kN/m3) σ'vp (kPa) Iεp,s Iεs Iεp,c Iεc Iε Iεmed hj (m) E (MPa) sj (m) 0.000 0.00 18 0.00 0.615 0.100 0.581 0.200 0.117 0.204 0.400 33.69 0.0003 0.400 0.02 18 3.28 0.615 0.120 0.581 0.207 0.135 0.220 1.300 40.35 0.0008 1.700 0.08 16.7 12.23 0.615 0.186 0.581 0.232 0.194 0.271 4.200 61.91 0.0020 5.900 0.29 18 46.63 0.615 0.399 0.581 0.310 0.384 0.320 1.000 82.30 0.0004 6.900 0.34 16.7 53.52 0.615 0.450 0.581 0.329 0.429 0.360 3.275 99.05 0.0013 10.175 0.50 18 80.34 0.615 0.615 0.581 0.390 0.577 0.406 1.695 118.54 0.0006 11.870 0.58 18 94.23 0.615 0.582 0.581 0.422 0.554 0.501 8.480 158.42 0.0029 20.35 1.00 18 163.68 0.615 0.410 0.581 0.580 0.440 0.484 20.350 271.44 0.0039 40.70 2.00 18 330.34 0.615 0.000 0.581 0.387 0.067 0.370 3.640 365.48 0.0004 44.34 2.18 18 360.15 0.615 -0.073 0.581 0.352 0.000 s (mm) 12.45 k (kN/m3) 8602.57 Foundation Design of a New Hospital Building in Kecskemét 106 Table A.3 – Schmertmann Method results for CPT2 z (m) z/B γ (kN/m3) σ'vp (kPa) Iεp,s Iεs Iεp,c Iεc Iε Iεmed hj (m) E (MPa) sj (m) 0.000 0.00 18 0.00 0.618 0.100 0.582 0.200 0.117 0.204 0.400 33.69 0.0003 0.400 0.02 18 3.28 0.618 0.120 0.582 0.208 0.135 0.217 1.000 39.18 0.0006 1.400 0.07 16.7 10.17 0.618 0.171 0.582 0.226 0.181 0.268 4.500 60.74 0.0021 5.900 0.29 18 47.02 0.618 0.400 0.582 0.311 0.385 0.343 3.500 92.10 0.0014 9.400 0.46 16.7 71.14 0.618 0.578 0.582 0.376 0.543 0.384 0.775 108.85 0.0003 10.175 0.50 18 77.48 0.618 0.618 0.582 0.391 0.578 0.407 1.695 118.54 0.0006 11.870 0.58 18 91.37 0.618 0.584 0.582 0.423 0.556 0.502 8.480 158.42 0.0029 20.35 1.00 18 160.82 0.618 0.412 0.582 0.581 0.441 0.484 20.350 271.44 0.0039 40.70 2.00 18 327.48 0.618 0.000 0.582 0.387 0.067 0.370 3.640 365.48 0.0004 44.34 2.18 18 357.29 0.618 -0.074 0.582 0.353 0.000 s (mm) 12.45 k (kN/m3) 8599.39 Table A.4 – Schmertmann Method results for CPT3 z (m) z/B γ (kN/m3) σ'vp (kPa) Iεp,s Iεs Iεp,c Iεc Iε Iεmed hj (m) E (MPa) sj (m) 0.000 0.00 18 0.00 0.615 0.100 0.581 0.200 0.117 0.206 0.600 34.47 0.0004 0.600 0.03 18 4.91 0.615 0.130 0.581 0.211 0.144 0.219 0.800 39.96 0.0005 1.400 0.07 16.7 10.43 0.615 0.171 0.581 0.226 0.180 0.308 8.775 77.49 0.0037 10.175 0.50 18 82.29 0.615 0.615 0.581 0.390 0.577 0.406 1.695 118.54 0.0006 11.870 0.58 18 96.18 0.615 0.582 0.581 0.422 0.554 0.501 8.480 158.42 0.0029 20.35 1.00 18 165.63 0.615 0.410 0.581 0.580 0.440 0.484 20.350 271.44 0.0039 40.70 2.00 18 332.29 0.615 0.000 0.581 0.387 0.067 0.370 3.640 365.48 0.0004 44.34 2.18 18 362.10 0.615 -0.073 0.581 0.352 0.000 s (mm) 12.45 k (kN/m3) 8662.14