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1 STUDY OF GROUTED DEEP FOUNDATIONS IN COHESIONLESS SOILS By SUDHEESH THIYYA KKANDI A DISSERTATION PRESENTED TO THE GRADUATE SCHOOL OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF DOCTO R OF PHILOSOPHY UNIVERSITY OF FLORIDA 2013
2 201 3 Sudheesh Thiyyak kandi
3 To my lovely wife Aswathy
4 ACKNOWLEDGEMENTS This dissertation would not have been possible without the help and encouragements o f many personalities around me. First and foremost, I would like to express my heartfelt gratitude to my advisor, Dr. Mich a el C. McVay for his excellent guidance, creative suggestions, and timely advices for the completion of this research. I have been ve ry fortunate to have an advisor like him. His dedication to research and teaching has always inspired me. He gave me the freedom to explore my own thoughts on the research, but provided the insightful discussions and guided me in the right direction whenev er I needed. I express my sincere thanks to other members of my doctoral committee, Dr. Dennis Hiltunen, Dr. Gary Consolazio, and Dr. Ajay Shanker for their valuable comments and suggestions. I also wish to convey my thanks to FDOT officials especially Mr. Peter Lai for providing the financial support during the period. I am deeply indebted to Coastal Engineering staff, Mr. Ryan Mackey, Mr. Jim Joiner, Mr. Victor Adams, and Mr. Richard Booze for the constant help and co operation rendered during the expe rimental works. I am thankful to Mr. Jose Hernando and his crew, State Materials Office, Gainesville, for all the helps provided to perform the field testing. I articulate my sincere thanks to Mr. Jon Sinnreich, Load T est Gainesville for providing me vario us monitoring instruments for my load tests. I would like to say thanks to all the members in geotechnical group for their help in various ways to bring out this work. I am grateful to my parents, parents in law, and grandparents in law for their love and encouragement. I am thankful to my brother, brothers in law, other relatives, and friends for t heir support and encouragement. I deeply thank my wife for her patience,
5 support, and encouragement during this journey and my son who has given me enormous ener gy and relieved stresses in my studies with his little smiles. Last but not the least I express my heartfelt thanks to Almighty God who made this work flawless and painless throughout the course of my doctoral research.
6 TABLE OF CONTENTS Page ACKNOWLEDGEMENTS ................................ ................................ ............................... 4 LIST OF TABLES ................................ ................................ ................................ ............ 9 LIST OF FIGURES ................................ ................................ ................................ ........ 11 ABSTRACT ................................ ................................ ................................ ................... 18 CHAPTER 1 INTRODUCTION ................................ ................................ ................................ .... 20 1.1 Problem Statement ................................ ................................ ........................... 20 1.2 Hypothesis ................................ ................................ ................................ ........ 23 1.3 Objectives ................................ ................................ ................................ ......... 23 1.4 Scope ................................ ................................ ................................ ................ 24 1.4.1 Analysis of the Previous Experimental Data on Jetted and Grouted Piles ................................ ................................ ................................ ....... 24 1.4.2 Numerical Modeling of Jetted and Grouted Piles ................................ ... 24 1.4.3 Develop a Design Methodology for Jetted and Grouted Piles ............... 25 1.4.4 Full Scale Field Installation and Testing of Single Jetted and Grouted Piles ................................ ................................ ................................ ....... 25 1.4.5 Small Scale Testing of Jetted and Groute d Pile Groups ........................ 26 1.4.6 Small Scale Testing of Post Grouted Drilled Shaft Groups .................... 26 1.4.7 Numerical Modeling of Post Grouted Dril led Shafts ............................... 27 1.4.8 Develop Axial Prediction Approach for Post Grouted Drilled Shafts ...... 27 1.4.9 Comparison of Side and Tip Groute d Versus Tip Only Grouted Foundations ................................ ................................ ........................... 28 1.5 Overview of Dissertation ................................ ................................ ................... 28 2 LITERATURE REVIEW ................................ ................................ .......................... 30 2.1 Pile Jetting ................................ ................................ ................................ ........ 30 2.2 Post Grouted Drilled Shafts ................................ ................................ .............. 31 2.3 Jetted and Grouted Precast Piles ................................ ................................ ..... 34 2.4 Cavity Expansion Theory ................................ ................................ .................. 35 2.5 Soil Pile Interaction ................................ ................................ ........................... 37 3 IND IVIDUAL RESPONSE OF JETTED AND GROUTED PILES ............................ 47 3.1 Analysis of the Previous Experimental Data on Jetted and Grouted Piles ........ 47 3. 1.1 Test Chamber and Instrumentation ................................ ....................... 47 3.1.2 Test Soil Properties and Test Chamber Soil Preparation ...................... 48 3.1.3 Residual Horizonta l Stress Around Pile ................................ ................. 50
7 3.2 Numerical Modeling of Jetted and Grouted Piles ................................ .............. 51 3.2.1 Material Models ................................ ................................ ..................... 51 3.2.2 Simulation of Grouting and Top Down Load Test ................................ .. 52 3.2.3 Lateral Stress Distribution and Lateral Soil Displacement During Pile Grouting ................................ ................................ ................................ 54 3.3 Design Methodology for Jetted and Grouted Piles ................................ ............ 56 3.3.1 Estimation of Jetted and Grouted Pile Grouting Pressures .................... 56 3.3.2 Estimation of Unit Skin Friction ................................ .............................. 57 3.3.3 Load Displacement Curve for Jetted and Grouted Piles ........................ 60 3.4 Full Scale Field Installation and Testing of Single Jetted and Grouted Piles .... 62 3.4.1 Soil Investigation at the Test Site ................................ ........................... 62 3.4.2 Design and Construction of Precast Piles Used for Jetted and Grouted Piles ................................ ................................ ......................... 64 188.8.131.52 Structural design ................................ ................................ ....... 64 184.108.40.206 Design and fabrication of grout delivery and jetting systems .... 64 220.127.116.11 Construction of precast piles and preparation for jetting ........... 65 18.104.22.168 Jetting of precast piles ................................ .............................. 67 22.214.171.124 Design and construction of concrete cap for jetted and grouted piles ................................ ................................ ............. 68 126.96.36.199 Side and tip grouting of the piles ................................ ............... 70 3.4.3 Measured Noise and Vibration During Pile Jetting and Grouting ........... 72 3 .4.3.1 Measured noise ................................ ................................ ........ 73 188.8.131.52 Measured ground surface vibration ................................ .......... 75 3.4.4 Axial Top Down Testing of the Jetted and Grouted Pil e ........................ 76 3.4.5 Construction and Testing of Comparison Drilled shaft ........................... 78 184.108.40.206 Construction of test drilled shaft ................................ ................ 78 220.127.116.11 Axial top down testing of drilled shaft ................................ ........ 79 3.4.6 Comparison of the Response of Jetted and Grouted Piles with Drilled Shaft ................................ ................................ ................................ ...... 79 3.4.7 Combined Torsion and Lateral Load Testing of the Jetted and Grouted Pile ................................ ................................ ........................... 80 18.104.22.168 Design and fabrication of Mast arm assembly .......................... 80 22.214.171.124 Test setup, instrumentation, and load test procedure ............... 81 126.96.36.199 Analysis of results ................................ ................................ ..... 82 4 GROUP RESPONSE OF JETTED AND GROUTED PILES ................................ 128 4.1 Group Testing of Jetted and Grouted Piles ................................ ..................... 129 4.1.1 Soil Preparation in the Test Chamber ................................ .................. 129 4.1.2 Design and Construction of Precast Piles ................................ ............ 130 4.1.3 Jetting of Precast Pil es for Each Group ................................ ............... 131 4.1.4 Side Grouting of the Piles ................................ ................................ .... 131 4.1.5 Static Top Down Test Prior to Tip Grouting ................................ ......... 132 4.1.6 Tip Grouting of the Piles ................................ ................................ ...... 132 4.1.7 Static Top Down Test After Tip Grouting ................................ ............. 133 4.1.8 Excavation of Jetted and Grouted Pile Groups ................................ .... 133 4.2 Analysis of Experimental Jetted and Grouted Pile Group Behavior ................ 133
8 4.3 Predicted Axial Response of Jetted and Grouted Pile Groups ........................ 137 4.3.1 Skin Resistance of the Groups ................................ ............................ 137 4.3.2 Load Dis placement Response of the Groups ................................ ... 138 5 GROUP RESPONSE OF POST GROUTED DRILLED SHAFTS ......................... 155 5.1 Group Testing of Post Grouted Drill ed Shafts ................................ ................. 156 5.1.1 Soil Preparation for Group Tests ................................ ......................... 156 5.1.2 Fabrication of the Reinforcing Cage and Tip Grout System for Dril led Shafts ................................ ................................ ................................ .. 156 5.1.3 Top Down Testing of Drilled Shaft Groups Prior to Tip Grouting ......... 157 5.1.4 Tip Grouting of Drilled Shafts ................................ ............................... 158 5.1.5 Top Down Testing on Tip Grouted Drilled Shaft Groups ...................... 159 5.1.6 Excavation of Tip Grouted Drilled Shaft Groups ................................ .. 159 5.2 Analysis of Experimental Post Grouted Drilled Shaft Group Behavior ............ 160 6 INDIVIDUAL RESPONSE OF POST GROUTED DRILLED SHAFTS .................. 179 6.1 Numerical Modeling of Post Grouted Drilled Shafts ................................ ........ 179 6.1.1 Material Models and Parameters ................................ ......................... 180 6.1.2 Modeling of Construction, Tip Grouting, and Top Down Load Tests .... 181 6.1.3 Skin Resistance of Tip Grouted Shafts ................................ ................ 183 6.1.4 Load Transfer Mechanism at Shaft Tip ................................ ................ 184 6.2 Develop Axial Prediction Approach for Post Grouted Drilled Shafts ............... 186 6.2.1 Estimation of Unit Tip Resistance vs. Tip Displacement ...................... 186 6.2.2 Estimation of Tip Area Increase Due to Grouting ................................ 189 6.2.3 Comparison of Prediction Approach with Field Test Data .................... 192 7 COMPARISON OF SIDE AND TIP GROUTED VERSUS TIP ONLY GROUTED FOUNDATIONS ................................ ................................ ................................ .... 207 7.1 Residual Horizontal Stress Around Deep Foundations ................................ ... 207 7.2 Maximum Tip Grout Pressure and Grout Bulb Formation ............................... 208 7.3 Axial Resistance ................................ ................................ ............................. 210 7.4 Group Interaction ................................ ................................ ............................ 210 8 CONCLUSIONS ................................ ................................ ................................ ... 215 LIST OF REFERENCES ................................ ................................ ............................. 220 BIOGRAPHICAL SKETCH ................................ ................................ .......................... 228
9 LIST OF TABLES Table Page 3 1 Measured lateral soil stress change due to jetting and grouting ......................... 85 3 2 Measured grout pressure and grout volume ................................ ....................... 85 3 3 Material properties used in PLAXIS ................................ ................................ .... 85 3 4 Measured and predicted grout pressures for 2.44 m long piles .......................... 86 3 5 Side resis tance prediction ................................ ................................ .................. 86 3 6 Estimation of required jet pipe diameter ................................ ............................. 87 3 7 Comparison of measured and predicted grout pressures ................................ ... 87 3 8 Comparison of the measured and predicted tip grout pressures ........................ 88 3 9 Limiting velocity suggested by AASHTO Designation R8 81 .............................. 88 3 10 Side resistance prediction for field jetted and grouted pile ................................ 88 3 11 Comparison of measured and predicted side resistanc e ................................ .... 89 3 12 Comparison of unit skin frictions for jetted and grouted pile vs. drilled shaft ...... 89 3 13 Forces and moments on the foun dation for the E7 T6 Mast Arm assembly (design wind speed = 130 mph) ................................ ................................ ......... 89 3 14 Dimensions of Mast arm assembly ................................ ................................ ..... 89 3 15 Forces and moments on the pile under maximum lateral load (54 kN) ............... 90 3 16 Comparison of mobilized and predicted (ultimate) torsional resistance .............. 90 4 1 Jetted and grouted piles grouting data ................................ ............................. 139 4 2 Highest horizontal soil stress increase near chamber boundary during grouting and group load test ................................ ................................ ............. 139 4 3 Side resistance prediction for single piles and groups ................................ ...... 140 4 4 Comparison of individual skin resistance prediction using the tip grout pressure and the pro posed method ................................ ................................ .. 140 4 5 List of parameters used for the group load displacement prediction ................ 141
10 5 1 Drilled shafts grouting data ................................ ................................ ............... 166 5 2 Increase of horizontal soil stress at shaft tip elevation during grouting & group test ................................ ................................ ................................ .................... 166 6 1 Material properties used in PLAXIS ................................ ................................ .. 195 6 2 Skin resistance of shafts before and after grouting ................................ ........... 195 6 3 Comparison of grouted and un grouted skin resistance from full scale tests in United States ................................ ................................ ................................ .... 196 6 4 Shafts and relevant parameters used for the prediction ................................ ... 197 6 5 Details of full sc ale field tests and relevant parameters ................................ .... 197 6 6 Shafts and relevant parameters used for the prediction ................................ ... 198 7 1 Comparison betwee n the measured tip grout pressures and spherical cavity expansion limit pressures ................................ ................................ ................. 213
11 LIST OF FIGURES Figure Page 2 1 Three zones during pile jettin g ................................ ................................ ............ 40 2 2 Pressurized tip grouting of drilled shafts ................................ ............................. 40 2 3 Grout distribution systems: Sleeve port and Flat jack types ............................... 41 2 4 Schematic of Jetted and grouted pile with grout delivery and jetting systems .... 42 2 5 Grout delivery systems for the top and bottom zones of pile .............................. 43 2 6 Jet nozzles and side grout membranes attached to piles ................................ ... 44 2 7 Excavated 0.406 m square x 6.1 m long jetted and grouted pile ........................ 45 2 8 Three zones around an expanding cavity ................................ ........................... 45 2 9 Soil pile interaction ................................ ................................ ............................. 46 3 1 (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ .......... 91 3 2 Earth pressure cells in the tes t chamber (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ................................ ..... 91 3 3 Grain size distribution of test soil ................................ ................................ ........ 92 3 4 Expansion pressur e vs. volume curves from Pressuremeter tests ..................... 92 3 5 Residual horizontal stress variation with time near grouted pile ......................... 93 3 6 Fini te element discretization ................................ ................................ ............... 93 3 7 FE mesh after the simulation of Jetted and grouted pile installation ................... 94 3 8 Stress distribution with radial distance during tip grouting (spherical cavity expansion) ................................ ................................ ................................ .......... 94 3 9 Radial displacement determined from the numerical simulation of side and tip grouting of 0.203 m square pile ................................ ................................ .......... 95 3 10 ................................ ......................... 96 3 11 Estimate of grout vertical stress coefficient, K g ................................ ................... 97 3 12 Comparison of load displacement curves for 0.152 m square pile ..................... 98
12 3 13 Comparison of load displacement curves for 0.203 m square pile ..................... 98 3 14 Comparison of load displacement curves for 0.406 m square pile ..................... 99 3 15 Layout of test piles and drilled shaft along with reacti on drilled shafts ................ 9 9 3 16 SPT blow count (N) profiles and the Unified Soil Classification (USC) at the location of test piles and drilled shaft ................................ ................................ 100 3 17 Typical grain size distributions for the silty sand at the site .............................. 101 3 18 Expansion pressure volume curves from Pressuremeter tests ......................... 101 3 19 Schematic diagram of jetted and grouted pile ................................ .................. 102 3 20 Grout delivery and jetting systems ................................ ................................ .... 103 3 21 Reinforcing cage with grout delivery and jetting systems (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ...... 103 3 22 Concrete placement for one of the precast piles (Photo courtesy of au thor, Sudheesh Thiyyakkandi) ................................ ................................ .................. 104 3 23 Preparation of pile for jetting (Photos courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ................................ ... 104 3 2 4 Pile jetting (Photo courtesy of author, Sudheesh Thiyyakkandi) ....................... 105 3 25 Schematic of precast concrete cap pile connection ................................ ....... 105 3 26 Longitudinal and cross section view of concrete cap with reinforcement details ................................ ................................ ................................ ............... 106 3 27 Placement and grouting of precast cap (Photos courtesy of author, James F Stephenson III) ................................ ................................ ................................ 107 3 28 Casting of cast in place concrete cap (Photo courtesy of author, James F Stephenson III) ................................ ................................ ................................ 107 3 29 Side grouting of piles (Phot o courtesy of author, Sudheesh Thiyyakkandi) ...... 108 3 30 Grout pressure volume response during side grouting of pile 1 ....................... 108 3 31 Gro ut pressure volume response during side grouting of pile 2 ....................... 109 3 32 Grout pressure volume response during tip grouting ................................ ........ 109 3 33 Grout pressure vs. pile head displacement during tip grouting ......................... 110
13 3 34 Instrumentation for noise and vibration measurement (Photos courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ...... 110 3 35 Location of construction equipments and vibration and noise monitors ............ 111 3 36 Noise measurement during pile jetting process ................................ ................ 112 3 37 Noise measurement during side grouting ................................ ......................... 112 3 38 Noise measurement during tip grouting ................................ ............................ 113 3 39 Peak particle velocity measurement during pile jetting ................................ ..... 113 3 40 Peak particle velocity measurement during grouting ................................ ........ 113 3 41 Axial top down test on jetted and grouted pile (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ .................. 114 3 42 Measured strain at different depths ................................ ................................ .. 115 3 43 Load displacement response of the jetted and grouted pile ............................. 115 3 44 Load distribution along the pile ................................ ................................ ......... 116 3 45 Lowering of reinforcing cage and concrete placement for test drilled shaft (Photos courtesy of author, Sudheesh Thiyyakkandi) ................................ ...... 117 3 46 Load distribution along the dr illed shaft ................................ ............................ 117 3 47 Load displacement response of drilled shaft ................................ .................... 118 3 48 Comparison of axial response of jetted and grouted pil e vs. drilled shaft ......... 118 3 49 Coordinate system used for representing forces and moments ........................ 119 3 50 Mast arm assembly fabricated fo r combined torsion and lateral load test (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ........ 119 3 51 Setting and bolting Mast arm assembly (Photos courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ .................. 120 3 52 Total Station for rotation and translation measurement (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ...... 121 3 53 String p ot arrangement (Photo courtesy of author, Sudheesh Thiyyakkandi) ... 122 3 54 Digital dial gage placement (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ................................ ... 123
14 3 55 Application of lateral load on the arm by pulling with a crane (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ .. 124 3 56 Tension load cell for the load measu rement (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ .................. 124 3 57 Torque vs. rotation response during combined torsion and lateral load test ..... 125 3 58 Lateral displacement components during the load test ................................ ..... 125 3 59 Lateral load vs. resultant lateral displacement during combined torsion and lateral load test ................................ ................................ ................................ 126 3 60 Torsional cracks and gaps after loading and unloading (Photos courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ...... 127 4 1 Soil compaction usin g vibratory plate compactor (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ .................. 142 4 2 Earth pressure cells placement (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ................................ ... 142 4 3 Reinforcing cages with grout delivery and jetting systems for group 1 piles(Photos courtesy of author, Sudheesh Thiyyakkandi) ............................... 143 4 4 Reinforcing cages with grout delivery and jetting systems for group 2 piles(Photos courtesy of author, Sudheesh Thiyyakkandi) ............................... 144 4 5 Attachment of semi rigid membranes and typical rubber nozzle used (Ph otos courtesy of author, Sudheesh Thiyyakkandi) ................................ .................... 145 4 6 Pile during jetting (Photo courtesy of author, Sudheesh Thiyyakkandi) ............ 146 4 7 Side grouting of pile in group 2 (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ................................ ... 146 4 8 Load test setup for jetted and grouted pile group (Photo courtesy of author, Sudheesh Thiyyakka ndi) ................................ ................................ .................. 147 4 9 Digital levels and invar staffs used for pile displacement monitoring (Photos courtesy of author, Sudheesh Thiyyakkandi) ................................ .................... 147 4 10 Jetted and grouted pile groups after excavation (Photos courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ .................. 148 4 11 Views of a tip grout bulb (group 1) (Photos courtesy of author, Sudhe esh Thiyyakkandi) ................................ ................................ ................................ ... 148 4 12 Load displacement response of group 1 prior to tip grouting ............................ 149
15 4 13 Load displacement response of gr oup 1 after tip grouting ................................ 149 4 14 Load displacement response of group 2 after tip grouting ................................ 150 4 15 Ground surface crack around group during axial loading (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ...... 150 4 16 Soil deformation profile (group 1 load test) ................................ ....................... 151 4 17 Soil deformation profile (group 2 load test) ................................ ....................... 151 4 18 Typical variation of horizontal stress around pile during axial load test ............ 152 4 19 Vertical stress variation below the center of the group footprint vs. beneath pile during the top down load test before tip grouting ................................ ....... 152 4 20 Typical vertical stress varia tion below the center of the group footprint vs. beneath pile during the top down load test after tip grouting ............................ 153 4 21 Comparison of group response before and after tip grouting of group 1 .......... 153 4 22 Predicted and measured load displacement response of group 1 .................. 154 4 23 Predicted and measured load displacement response of group 2 .................. 154 5 1 PVC casing positioned before filling the test chamber (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ...... 167 5 2 Test chamber in fully filled state (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ................................ ... 167 5 3 Reinforcing cage and grout distribution system (Photos courtesy of author, Sudheesh Thiyyakkand i) ................................ ................................ .................. 168 5 4 Pulling the casing out (Photos courtesy of author, Sudheesh Thiyyakkandi) ... 169 5 5 Group load test setup (Photo courte sy of author, Sudheesh Thiyyakkandi) ..... 169 5 6 Tip grouting (Photo courtesy of author, Sudheesh Thiyyakkandi) .................... 170 5 7 Setup for in dividual shaft loading (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ................................ ... 170 5 8 Excavated group 1 shafts (Photos courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ................................ ... 171 5 9 Excavated group 2 shafts (Photo courtesy of author, Sudheesh Thiyyakkandi) ................................ ................................ ................................ ... 172
16 5 10 Load displacement response of group 1 shafts ................................ ................ 172 5 11 Load displacement response of group 2 shafts ................................ ................ 173 5 12 Load displacement response of group 1 shafts during the group and individual load tests. ................................ ................................ ......................... 173 5 13 Load displacement response of group 2 shafts during the group and individual load tests ................................ ................................ .......................... 174 5 14 Displacements of all t he shafts in group 1 during south shaft loading .............. 174 5 15 Displacements of all the shafts in group 2 during the individual shaft loading .. 175 5 16 Vertical stress measurement beneath group footprint during the group load tests ................................ ................................ ................................ .................. 176 5 17 Vertical stress measurement beneath group footprint during the individual shaft loadin g ................................ ................................ ................................ ..... 177 5 18 Variation of residual horizontal stress after different stages measured using earth pressure cells at 15 cm (0.5 ft) away from shafts ................................ .... 178 6 1 Typical Finite element discretization ................................ ................................ 199 6 2 displacement during axial load test (Group1 East shaft) ................................ .. 200 6 3 Unit tip resistance versus tip displacement for un grouted and grouted shafts from FEM ................................ ................................ ................................ .......... 200 6 4 Mechanism of load transfer during axial loading of base grouted shaft ............ 201 6 5 Values of recommended by various investigators and from load tests .......... 202 6 6 Predicted and measured response for Georgia Tech load test ........................ 202 6 7 Conceptual normalized tip resistance displacement plot ................................ .. 203 6 8 Comparison of unit tip resistance displacement response (2.44 m long shaft) 203 6 9 Comparison of unit tip resistance displacement response (3.96 m long shaft) 204 6 10 Determination of mobilized tip stress, p s from log log plot of tip stress (embedme nt strain gage) vs displacement (shaft: S1 FJ1) .............................. 204 6 11 Area ratio ( A r ) versus NGP / ................................ ................................ .... 205 6 12 Predicted and measured tip loa d displacement response (Clearwater, FL ) ..... 205
17 6 13 Predicted and measured tip load displacement response (Houston) ............... 206 6 14 Predi cted and measured response (Audubon Bridge project, Louisiana) ......... 206 7 1 Group behavior of jetted and grouted piles and post grouted drilled shafts ...... 214
18 Abstra ct of Dissertation Presented to the Graduate School of the University of Florida in Partial Fulfillment of the Requirements for the Degree of Doctor of Philosophy STUDY OF GROUTED DEEP FOUNDATIONS IN COHESIONLESS SOILS By Sudheesh Thiyyakk andi May 2013 Chair: Michael. C. McVay Major: Civil Engineering Grouting of deep foundation subsequent to installation has become popular due to its eff ectiveness in improving axial capacity under serviceable displacement P ost tip grouting of drilled shafts has been successfully employed worldwide to mobilize a significant portion of tip resistance under small displacements. Recently, Florida Department of Transportation developed a new jetted and grouted precast pile system and the construction of the pile utilizes the advantages of several proven deep foundation installation techniques This research focused on the individual and group behavior of jetted and grouted piles and tip grouted drilled shafts in cohesionless soils. Experimental study and nu merical modeling of both types of foundations were performed to investigate the response in individual and group scenario. Based on th e study, a methodology for jetted and grouted piles that predicts expected grout pressures during grouting, unit side and tip resistance and the load displacement response of the pile is proposed. In case of post grouted drilled shafts, t he study found that the increased axial capacity under service able displacements depended mainly on preloading effects and the increased tip area provided by the grouting process. A simple prediction approach for estimating the
19 tip capacity of grouted shafts utilizing cone penetration resistance was suggested based on the results of the study. The validity of the proposed approach was verified by the analysis of full scale case studies of grouted shafts reported in the literature. The experimental group study at 3 x pile/shaft diameter (D) center to center spacing revealed that the jet ted and grouted piles behaved as a block during axial loadi ng whereas the post grouted drilled shafts acted indepen dently of one another, i.e., n egligible group interaction It was also identified that the side grouting of a foundation prior to tip grouting has a substantial influence on improving the axial capa city. T he side grouting was found to significantly increase the grout pressure developed during tip grouting and helps in the formation of tip grout bulb by spherical cavity expansion process and thus improves the unit tip resistance of the foundation.
20 CHAPTER 1 INTRODUCTION 1.1 Problem Statement Deep foundations are widely used to support buildings, bridges, signage and other structures for the transfer of superstructure loads to soil under acceptable vertical and lateral displacements. In the past, foundat ions of choice were driven concrete piles However, noise and vibration from dynamic pile driving is a critical issue in urban environments (Selby 1991; Woods 1997; White et al. 2002; Svinkin 2006). Other alternatives are the use of drilled shafts and continuous flight auger (CFA) piles due to 2007). Unfortunately, a significant portion of a dri due to the vertical displacement required for mobilization. Displacements of about 10 15% of shaft diameter are required to fully mobilize the end bearing, whereas skin resistance fully develops at a displacement of about 0.5 1% of shaft diameter (Bruce 1986; Mullins and Dapp 2006). Current service design (e.g., AASHTO 2010) limits the vertical displacements of bridge substructure components to less than 50 mm, which significantly limits the mobilized tip resistance To regain some of the unused tip capacity, post grouting the drilled shaft tip has been successfully employed worldwide over the last five decades. The post grouting has been used in Asia and Europe to improve pile capacity since the early 1960s (Bologne si and Moretto 1973; Gouvenot and Gabix 1975; Stocker 1983). The post grouting process consists of the following: (1) casting a drilled shaft with a grout delivery system integrated to the rebar cage; and (2) injecting high pressure colloidal grout
21 beneath shaft base after sufficient shaft curing, which preloads the in situ soil, and f ills in all voids and anomalies in the vicinity of the shaft tip. Additionally tip grouting provides a proof test for every shaft, resu lting in higher LRFD (Load Resistance F actor Design) resistance factors, (Mullins et al. 2006). Several case studies (Mullins et al. 2001, 2004 and 2006; Ruiz et al. 2005; Duan and Kulhawy 2009; Youn and Tonon 2010; Dapp and Brown 2010, Dai et al 2010) have been performed over the last decad e to identify the effectiveness of tip grouting, the factors influencing the improvement of axial capacity, and the fundamental mechanisms involved. A number of design methodologies have been proposed for post grouted drilled shafts. For example, Mullins e t al. (2006) has presented a design methodology based on Federal Highway grouted shaft using Standard Penetration test (SPT) blow count (N) to estimate tip grout pressures and subse quent mobilization of tip resistance. Ruiz (2005) has proposed a design approach called Axial Capacity Multiplier (ACM) based on nine case histories, that uses CPT tip resistance, q c The method considers three factors: (1) soil compression under pile tip, (2) enlarged tip area with grout tip bulb formation, and (3) side shear reversal, as the major contributors to the improved capacity of tip grouted shafts. However, neither approach identifies the influence of tip grouting on the ultimate end bearing of a post grouted drilled shaft. Although pressurized tip grouting improves the tip resistance of drilled shafts under serviceable displacement, still remaining was the issue of construction quality control (i.e., the structural integrity of the shaft) and th e fact that they have lower skin friction than driven piles inherent in the installation process (Jalinoos et al. 2005, Kog
22 2009, Meyerhof 1976). Recently, Florida Department of Transportation (FDOT) developed a new jetted and grouted precast pile system ( with side membranes) in cohesionless soils and the construction of the new pile combines the advantages of several proven deep foundation installation techniques (McVay et al. 2009: FDOT report BD545 31): (1) use pre cast reinforced concrete pile to elimin ate the unknown quality of the cast in situ drilled shaft/CFA pile; (2) jetting the precast pile which is expected to minimize the construction noise and vibration (Tsinker 1988); and (3) grouting the side and tip to maximize skin and tip resistance. The n ew pile which is jetted and al. 2012). An appropriate design methodology to predict expected grout pressures during grouting, unit side and tip resistance, and load d isplacement response need to be developed for the new pile. In addition, the constructability and applicability of the pile under typical field conditions has to be validated by performing full scale field installation and testing. FDOT UF research project (BD545 31) found from the full scale torsion testing of the new pile in a larger test chamber that the pile has large torsional capacity and suggested that the pile may be used as the foundation for Mast arm structures (i.e., the structures supporting hig hway signs and traffic signals). However the response of the new pile needs to be verified in typical field condition by performing the full scale testing. In case of the group placement of piles or shafts, the spacing of piles/shafts within a group is ge nerally a tradeoff; at possible minimum spacing, which reduces the high cost of the reinforced concrete caps as well as the group interference. When the piles or shafts are too close, the axial capacity of the group is significantly reduced, i.e.,
23 the axia l group resistance may be significantly less than the sum of the individual pile (or shaft) resistance (group efficiency factor < 1). A group efficiency factor of one (1.0) identifies that the stresses transferred to soil from each individual pile does not overlap with adjacent piles. Past research has shown that a center to center (c/c) spacing of three times the pile or shaft diameter (3D) will result in a group efficiency factor of one (1) for both the driven pile and drilled shaft groups. However, the b ehavior of post grouted drilled shafts and jetted and grouted precast piles in group placements are currently unknown. This dissertation focus es on the individual and group behavior of the aforementioned grouted deep foundations: (1) jett ed and grouted pre cast piles and (2) post grouted drilled shafts in cohesionless soils 1.2 Hypothesis Individual and group response (both skin and tip) of the piles or drilled shafts subjected to side and tip grouting will be significantly different from that of piles/shaft s undergone tip grouting only, because of the difference in the mechanism of grout bulb formation (e.g., cavity expansion process). 1.3 Objectives Specific o bjectives of this research include the following: Develop a design methodology to predict anticipated g rout pressures during grouting, unit side and tip resistance, and load displacement response of the jetted and grouted precast piles in cohesionless soils. Validate the constructability of the jetted and grouted precast piles in typical Florida sand. C ompare the axial, and the combined torsional and lateral response of the jetted and grouted precast piles with that of similar sized drilled shaft
24 Investigate the group interaction of the jetted and grouted precast piles at typical 3D spacing in cohesio nless soils. Investigate the group behavior of post grouted drilled shafts at typical 3D spacing in cohesionless soils Study the grout flow pattern and associated bulb formation during the tip grouting of drilled shafts, and the load transfer mechanism at shaft tip during the subsequent axial loading. Develop a prediction approach for the tip resistance of base grouted drilled shafts utilizing the cone penetration resistance (CPT, q c ). C ompar ison of the responses of side and tip grouted foundations v ersus tip only grouted foundations. 1.4 Scope 1.4.1 Ana lys is of the P revious Experimental D ata on J etted and Grouted P iles A detailed analysis of the previous experimental results on j etted and grouted precast piles reported by McVay et al. (2009) was performed Specifically the soil stress change (both horizontal and vertical stress) in the vicinity of the piles during the installation (i.e., jetting and grouting ) and static top down test were investigated In addition the variation of horizontal stress around the piles with the passage of time subsequent to the installation was also analyzed. 1.4.2 Numerical Modeling of J etted and Grouted P iles Numerical m odeling of the experimental j etted and grouted piles (McVay et al. 2009) were carried out to investigate the s oi l stresses around the piles using the two dimensional finite element package, PLAXIS 2D, developed by PLAXIS b. v., Delft, Netherlands The results obtained from the numerical analysis were compared with the experimental results The findings from the fin ite element analysis were also used for developing a design methodology for the pile.
25 1.4.3 Develop a D esign M ethodology for Jetted and Grouted Pile s Based on the experimental and numerical analys e s a design methodology for the j etted and grouted piles was pr oposed. Specifically, the approaches to predict the expected grout pressures during grouting, unit skin friction and the load displacement response of the pile in cohesionless soil were suggested The predicted responses of the pile using the proposed meth odology were compared with experimental and numerical response s and were found to match quite well. 1.4.4 Full S cale F ield I nstallation and T esting of Single Jetted and Grouted Piles Installation and testing of two 0.71m square x 5.5 m long (28 in square x18ft l ong) j FL in connection with FDOT UF ongoing research project ( BDK 75 977 41 ). A detailed soil exploration was performed at the test site, which include d Standard Pe netration Tes ts (SPT), Cone Penetration T ests (CPT), Pressurem eter T est s (PMT), and Dilatometer Tests (DMT). The soil at the test site was predominantly silty sand (SM). The installation of the piles validated the constructability of the Jetted and grouted pile in typi cal Florida field condition. Noise and ground surface vibration monitoring was also carried out during the jetting and grouting of the piles, and it was found that the pile is well suited for urban areas, where the construction noise and vibration are crit ical concern s Static top down test and combined torsion and lateral load test were performed on the piles T he axial response of the pile w as compared with that of similar sized drilled shafts installed at the same test site The combined torsion and late ral load test on the pile was performed by attaching a full scale Mast arm assembly to the top of pile and subsequently a pplying lateral load in increments at a standoff distance
26 (eccentricity) of 10.67 m (35 ft) from the axis of pile Rotation and transla tion of the pile were measured during the test. 1.4.5 Small Scale Testing of Jet ted and G routed Pile Groups Small scale testing of two j et ted and grouted pile groups was conducted to study the soil structure interaction between the piles within the group. The se tests were performed as a part of the experimental study involved in FDOT UF research project: BDK 75 977 07 The study was limited to four piles in each group with the typical 3D center to center spacing (i.e., three times the precast pile width/diameter) between the piles. Coastal Engineering Lab, University of Florida T he d isplacement of individual pile s, the deformation of soil, and the soil stress es within and outside the groups were monitored during the group tests The measured results were used to identify the behavior of groups under top down loading. An approach to predict the axial group response of the piles (@3D spacing) was also suggested and a reasonable agreement was found between the measured and predicted responses. 1.4.6 Small Scale Testing of Post Grouted Drilled Shaft Groups Two small scale group tests of grout tipped drilled shafts were conducted in the test chamber to investigate the group behavior under axial lo ading The study was limited to the groups with four shafts at 3D center center spacing The diameter of the shafts in both groups was same ( 0.216 m ), but different embedment depths; 2.44 m for group 1 ( i.e., Length/Diameter ~11) and 3.96 m for group 2 (i. e., L/D ~18). Note that t he smaller diameter shafts were selected to minimize the chamber boundary effects The goal of the first group test was to study the factors inf luencing the axial capacity of post grouted drilled shaft, the grout flow pattern, and the group behavior at typical 3D
27 spacing Whereas t he objective of the second test was to validate the results of the first group tests for greater embedment depths and investigate the feasibility of staged grouting to improve the axial capacity of grout t ipped shafts The groups testing as well as individual shaft tests were performed to estimate group interaction Measured axial top down testing data included soil defor mation in the vicinity of shaft, load displacement response of individual shafts, verti cal and horizontal soil stresses alongside and beneath individual shafts and the group These measured data were then used to estimate the interaction between the shafts within each group during the axial loading. 1.4.7 Numerical Modeling of Post Grouted Drille d Shafts To investigate the load transfer mechanism at the shaft tip, finite element analysis of p ost grouted drilled shafts was conducted using PLAXIS 2D. The drilled shaft and soil within the test chamber was simulated by an axisymmetric model. The const ruction of shaft, tip grouting, and top down load test w as modeled. Un grouted drilled shafts were also modeled for comparing the mobilized tip resistance during top down load test. The load transfer mechanism captured from the numerical analysis was later used for dev eloping a prediction methodology for the tip resistance of post grouted drilled shafts. 1.4.8 De velop A xial P rediction A pproach for Post G routed D rilled S hafts A prediction method for the tip capacity of post grouted drilled shaft utilizing the con e penetration resistance was developed based on the experimental and numerical analys e s. Specifically, an approach to predict the unit tip resistance displacement response and an equation to compute the final tip area of grouted shaft were suggested. The e xpression to estimate the final tip area was obtained from the
28 regression analysis of full scale field test data with known CPT value s The suggested prediction method was validated by applying to some of the full scale field tests available in literature. 1.4.9 Comp arison of S ide and T ip G routed V ersus T ip O nly G routed F oundations The response of side and tip grouted deep foundation was compared with tip only grouted foundations in both individual and group placements based on the results of experimental and nu merical studies. The influence of side grouting of a pile /shaft prior to tip grouting on the soil stress state around pile/shaft, maximum tip grout pressures and tip grout bulb formation was analyzed. T he difference in the interaction of piles/shafts at typical 3D spacing was also discussed. 1.5 Overview of Dissertation An overview of the following chapters follows. Chapter 2 is a literature review and provides an overview of pile jetting, pile grouting, past research into post grouted drilled shafts, jette d and grouted piles, cavity expansion theory, and soil pile interaction. Chapter 3 presents the individual response of j etted and grouted piles in cohesionless soils. An analysis of soils s tress measurements near the piles from the previous experimenta l st udy, numerical modeling of j etted and grouted pile s and a design methodology for the pile s are described in detail s. Chapter 4 describes the group behavior of jetted and grouted piles at typical 3D spacing based on the group tests. Soil preparation, const ruction of precast piles, installation of the piles (i.e., jetting and grouting), group load tests, analysis of the test results, and identified group interaction are discussed comprehensively.
29 Chapter 5 presents the group testing of p ost grouted drilled shafts at 3D spacing and the axial group efficiency of the shafts based on the measured soil stress, pile and soil deformation during the group and individual loading of the shafts. The group response of p ost grouted drilled shafts was also compared with t hat of j etted and grouted pile groups described in C hapter 4. Chapter 6 demonst rates the numerical modeling of individual p ost grouted drilled shafts and the axial prediction approach for the grouted shafts developed based on the experimental and numerical study Chapter 7 presents a comparison of the effectiveness of side and tip grouting versus tip only grouting on improving the capacity of deep foundations in cohesionless soils.
30 CHAPTER 2 LITERATURE REVIEW This C hapter reviews past studies on pile jetting p os t grouted drilled shafts, and j etted and grouted piles. It also reviews past research on cavity expansion theory and its geotechnical applications and soil structure interaction of deep foundations. 2.1 Pile J etting J etting of piles utilizing pressurized wate r has been widely used to aid pile penetration into dense to very dense sand layers to expedite pile driving and minimize vibration (Tsinker 1988, Gunaratne et al. 1999, Gabr et al. 2004). Jetting can assist the pile installation in several ways: (1) the j etting pressure may loosen (erode) the soil at the tip of the pile; (2) jetting may increase local pore water pressure and hence decrease effective stress, which eases pile penetration; (3) the upward flow of the jetting fluid lubricates the pile and assis ts its downward movement (Tsinker 1988). In 1959, Shestopal developed the following flow rate equation to estimate the water requirements for jetting into sandy soil: (2 1) w here Q = flow rate (m 3 /hr ) D = pile diameter or width (m) d 50 = av era g e size of sand particles (mm) l = desired submerged length of pile (m) C = 0.1 for dry sand and 0.017 for saturated sand stratum k = ( n l n ) / l = av erage permeability coefficient (m/ day)
31 Tsinker (1988) has identified three zones in the jet hole structure during pile jetting in sand stratum as shown in Figure 2 1 : (1 ) sand water mixture immediately beneath the pile tip (zone 1), (2 ) excess water pumped into zone 1 escapes to the surface alongside the pile (zone 2) and (3) sand water mixture around zone 2 at high pore pressures (zone 3). This excess pore pressure dissipates immediately after jetting in sand. Gunaratne et al. (1999) found that the lateral load capacity of a jetted pile is significantly less than that of driven pile due to the soil disturbance resulting from the jetting process. Gabr et al. (2004) has identified that pile insertion rate increases with increase in flow velocity for a given flow rate. Recently Giken Seisaku sho Ltd. has pipe piles (White et al. 2002). In soft soils, the pile is pushed, whereas in dense, stiff or hard soils, a disposable jet tip is attached to assist in the pile installation by jetting. 2.2 Post G routed D rilled S haft s The first known published test results using shaft grouting were Gouvenot and Gabix (1975). Their results indicated an increase in shaft fric tion of about 250% over un grouted bored piles. Bruce (1986) has presented a review of published work on pile construction and the benefit of post grouting between 1975 and 1985. More recently, tip and shaft grouting were used for piles and drilled shafts in sands (Plumbridge and Hill 2001). A number of different apparatus for side grouting (e.g., Joer et al. 1998 McVay et al. 2009 ) and tip grouting (Mullins et al. 2001) have been developed. Typical grout mixes used for grouting drilled shaft tips are ceme nt, sand, and water. Micro fine materials (e.g., fly ash, bentonite, etc.) are also used to partially replace cement and improve pumpability.
32 As mentioned earlier, post grouting drilled shaft tips has become popular worldwide due to its effectiveness in mo bilizing a large portion of available tip resistance under service displacement s The grouting of a shaft base by injecting a high pressure grout fills in any anomalies present beneath the shaft tip and pre stresses the underlying soil (Figure 2 2) T hree types of grout distribution systems are commonly used in practice to deliver the grout at tip : (1) stem type (2) sleeve port type, and (3) flat jack type (Mullins et al. 2001). The stem type is the simplest form of grout distribution system and consists o f a pipe end at the shaft tip. This is not an efficient grout distribution system, and hence, only utilized in the remediation of substandard shafts with inadequate capacity ( Mullins et al. 2001) The sleeve port type, also known as a tube a Manchette, pri marily consists of a pipe network at shaft tip with pre drilled holes and the pipe network is connected to grout tubes at the top of the shaft (Figure 2 3) The grout system has both grout entry and exit pipes, which allow s the flushing of the grout system and makes the re grouting of the shaft possible if necessary. The pipe network at the shaft tip is wrapped with a rubber membrane at the location of the holes to prevent blockage at the hole during the casting of the shaft and allows the grout to flow out during the grouting stage. This also prevents the return of pumped grout to the pipe network. Flat jack type consists of stem type pipes ends at the shaft tip within a plate and membrane system ( Figure 2 3 ) This plate and membrane system confine s the gro ut mass and prevents mixing with the surrounding soil (Mullins et al. 2001). The drawback of the flat jack type apparatus is that the grout lines cannot be flushed properly when re grouting is necessary unlike the sleeve port type grout distribution syste m
33 The effectiveness of tip grouting on improving the axial capacity of drilled shafts was investigated by several researchers (Mullins et al. 2006; Ruiz 2005; Duan and Kulhawy 2009; Youn and Tonon 2010; Dapp and Brown 2010 ; Dai et al 2010) over the last decade Mullins et al. (2006) ha ve proposed a prediction approach for the unit tip resistance of p ost grouted drilled shaft based on the regression analysis of a number of full scale field tests. Following are the steps involved in the prediction approach suggested by Mullins et al. (2006): Estimate the unit end bearing ( q b ) at a displacement of 5% of diameter of drilled q b =0.057N in MPa; N=uncorrected SPT blow count) Estimate the ultimate side resistance ( F s ) of the shaft. Determine the maximum expected grout pressure ( GP max ) by dividing the ultimate side resistance ( F s ) the cross sectional area ( A ). Convert the maximum expected grout pressure ( GP max ) to a dimensionless quantity, called Grout Pres sure Index ( GPI ) by dividing with the ultimate unit end bearing ( q b ). Determine the Tip Capacity Multiplier (TCM) using Equation 2 2. (2 2) w here %D = displacement expressed as the ratio of shaft di ameter Estimate the grouted unit end bearing as the product of TCM and the ultimate un grouted unit end bearing ( q grouted =TCM. q b ) Ruiz (2005) has developed design charts for the total capacity of p ost grouted drilled shaft s based on the eight reporte d case studies. The method called Axial Capacity Multiplier (ACM) uses the maximum anticipated grout pressure, and the Davisson failure load for equivalent un grouted shaft as the input parameter s to predict
34 the axial capacity of a Post grouted drilled s haft for a given pile head displacement Youn and Tonon ( 2010 ) reported that the Axial Capacity Multiplier (ACM) approach significantly over predicted the total resistance of grouted drilled shaft in a case study at the Brazo River Bridge, TX 2.3 Jetted and G routed P recast P ile s The construction of j etted and grouted precast pile s is comprised of four distinct phases : (1) construction of precast pile with jetting and grout distribution system s (2) p ressurized water jetting of the pile into ground, (3) s ide gr outing of the pile, and (4) tip grouting (McVay et al. 2009 ; Thiyyakkandi et al. 2012 ). Figure 2 4 shows the schematic of the j etted and grouted pile with grout delivery and jetting systems. The pile consists of separate grout delivery pipes for side and t ip grouting. The side grout system is separated into a top and bottom grouting zones with their own pipe network, Fig ure 2 4 Each of the side grout pipes has an entry and an exit outlet to allow staged/repeated grouting (Figure 2 5 ). To allow repeated gro uting, the bottom half of each grout pipe has a series of holes drilled into them with Gum rubber covering membranes Fig ure 2 5 A center jetting pipe is used to provide pressurized water at tip for pile jetting In the case of large size ( width ) piles, t he jet pipe branches off to four or five pipes at bottom for the uniform distribution of water at tip. The nozzle at the end of the jet pipe (Fig ure 2 4 and 2 6 ) not only increases the water velocity but also minimizes the water consumption during jetting The nozzle also prevent s sand or fines from ingression into the jet pipe after jetting, which can result in grout blockage as the jetting pipe and nozzle are later used for the tip grouting. pumping, a grout mix consisting of cement, micro fine fly ash and water was used (McVay et al. 2009 ;
35 Thiyyakkandi et al. 2012 ). To prevent grout flowing along the weakest path during side grouting, membranes were attached to the pile, ( Figure 2 4 and 2 6) These membranes confine the grout zones and improve radial expansion during grouting, resulting in major principal stresses in the horizontal/radial direction near the pile. The membranes also prevent the mixing of grout with the soil which improves bondi ng between the grout and the pile (McVay et al. 2009 ; Thiyyakkandi et al. 2012) Previous study shows that the piles possess very high axial and torsional resistances (McVay et al. 2009; Lai et al. 2010) Thiyyakkandi et al. (2012) have reported that the u nit skin friction of a Jetted and grouted pile is about 5 times that of similar sized drilled shafts Figure 2 7 displays the 0.406 m square x 6.1 m long jetted and grouted pile after excavation (McVay et al. 2009; Lai et al. 2010). Excavation of all the t est piles revealed that the side grout bulb had surround ed the entire perimeter of the pile s at each grout zone and was well bond ed to the pile s ( McVay et al. 2009 ; Thiyyakkandi et al. 2012) 2.4 Cavity Expansion Theory C avity expansion analysis offers useful solutions to a variety of problems in geotechnical engi neer ing, including in situ testing such as pressure meter and cone penetration testing, pile driving, pile loading to failure, tunnel deformation, and finally, the process of grouting a pile in situ. Initially, cavity expansion theory focused on solving metal indentation problems (Bishop et al. 1945; Hill 1950) The cavity expansion theory was first applied in the geotechnical engineering field by Gibson and Anderson (1961) for the interpretation of pr essuremeter tests The theory has been progressively refined and applied to various geotechnical problems in the last four decades (Palmer 1972; Vesic 1972; Hughes et al. 1977; Carter et al. 1986; Yu and Houlsby 1991; Salgado and Randolph 2001; Salgado and Prezzi 2007). Yu (2000) has presented fundamental
36 solutions for the cavity expansion problems, major developments, and applications in the field of geotechnical engineering. Cavity expansion process es are of two basic types: (1) e xpansion from a finite r adius and (2) expa nsion from zero initial radius; i e. cavity creation problem (Salgado et al. 1997 ) In the first case, an ever increasing pressure is required for continuing the expansion. The cavity wall pressure approaches the limit pressure (i.e., s teady expansion pressure) only when the cavity radius approaches infinity (Salgado et al. 1997). However in the second case, the cavity radius is initially zero and hence expansion to a finite radius (i.e., cavity creation) would be sufficient to develop t he limit pressure and further expansion occurs under constant cavity pressure. This is due to the fact that the expan sion from a zero radius to a finite cavity radius is equivalent to the expansion of a n initially existing cavity to an infinit e radius (Sal gado et al. 1997). Based on the strain level s the surrounding region of an expanding cavity can be characterized by three distinct zones: (1) Plastic zone, (2) nonlinear elastic zone, and (3) linear elastic zone, as depicted in Figure 2 8 (Salgado et al. 1997) In the plastic zone, the material has already failed due to the large stress state. In the nonlinear elastic zone, the material has yielded, but not failed because the stresses are not enough to cause the failure; and in the linear elastic zone, the stress strain response is within the elastic limit. In the case of jet ted and grouted pile s side grouting resembles the expansion of a cylindrical cavity from a finite radius and tip grouting resembles the expansion of a spherical cavity Hence, the cav ity expansion solutions may be utilized in developing the design methodology for the pile. In this research the elastic perfectly plastic closed form solutions of Yu and Houlsby (1991) and the limit pressure charts for cylindrical and
37 spherical cavities g iven by Salgado and Randolph (2001) we re used Yu and clos ed form solution are based on an elastic perfectly plastic soil with Mohr Coulomb failure criterion and a constant rate of dilatation ( ) Yu and Houlsby (1991) have presented a straightforward procedure for constructing the pressure expansion curve s and calculating limit pressure s for expanding cylindrical and spherical cavities. Salgado and Randolph (2001) presented a numerical method f or solution of cavity expansion problems taking into account stress equilibrium and strength and flow assump tions which resulted in charts for cylindrical /spherical cavi ty expansion limit pressures ( p lim ) as a function of soil strength ( c = critical sta te friction an gle), relative density ( D r ), and depth or initial lateral /mean in situ stress for sands 2.5 Soil P ile Interaction Most deep foundations consist of a group of piles or drilled shafts The piles or shafts are placed at the minimum possible spaci ng to reduce the cost of the concrete pile/shaft cap Failure of the group may occur either by failure of the individual piles or failure as an overall block The load capacity of a group of vertically loaded piles /shafts can, in many cases, be considerabl y less than the sum of the capacities of the individual piles / drilled shafts comprising the group as there will be shear transfer occurring through the soil from one pile/shaft to other Generally, a group efficiency factor of one (1) means that the shear stress transfer from one pile/shaft is not overlapped with that of an adjacent pile/shaft Past research has shown that a group efficiency factor of one (1) is achieved at a minimum center to center spacing of three times the diameter of the pile/shaft
38 T he pile/shaft soil interaction may be characterized as in Figure 2 9 If one considers the case of ultimate pile/shaft capacity, maximum side shear stress ( o ) is mobilized along the surface of the pile For any vertical slice (Figure 2 9 ), the shear stress ( 1 2 ) must always diminish with radius r and is negligible at a radial distance r m (radius of influence) Hence, it is evident that any pile placed wi thin the distance r m of an adjoining pile, undergoes shear transfer and settlement from the loaded adjoining pile, without any load being applied to the pile. In the case of a side grouted pile, pressure grouting increases both horizontal h ) and s hear strength of the soil around the pile shear strength, the shear modulus ( G ) also increases Consequently, for any applied load, the soil shear strain z r ) must be smaller Hence, in the case of ultimate capacity, muc h larger side shear stresses are expected alongside the grouted pile/shaft perimeter At the radial distance r m the shear stress is much greater for grouted pile s compared to conventional cast in situ piles/ shafts B ut shearing strain is smaller due to high shear m odulus for any radial distance, r compared to a non grouted cast in situ pile/shaft This suggests that a group of grouted piles/shaft could have greatly reduced efficiency factors for typical spacing, e.g., 3 D A low shear strain is expected w ithin the footprint of the group due to the increased confining stress and shear modulus, and much higher shear strain is expected outside the footprint where shear modulus is greatly diminished Consequently, the grouted pile group may fail through block failure Hence group efficien cies of the jetted and grouted pile group s are expected to be less than one at typical spacing of 3D But for a tip grouted shaft (no side grouting) group, a
39 combined conventional single shaft summation for side shear is expec ted, however, the tip resistance may exhibit a group footprint.
40 Figure 2 1 Three zones during pile jetting Figure 2 2. Pressurized tip grouting of drilled shafts Pumped grout Pre stress the soil Negative skin friction mobilization Zone 1 Zone 3 Zone 2
41 Figure 2 3. Grout distribution systems: Sleeve port and Flat jack type s Sleeve port type Flat jack type (Source: FD OT Report BC165 v1)
42 Figure 2 4. Schematic of Jetted and grouted pile with grout delivery and jetting systems Side grout system Jet/tip grout pipe Jet Nozzle Membrane Precast pile Grout orifice (G um rubber membrane)
43 (Source: FDOT Report BD545) Figure 2 5. Grout delivery systems for the top and bottom zones of pile
44 ( Source: FDOT Report BD545 ) Figure 2 6. Jet nozzles and side grout membranes attached to piles
45 ( Source: FDOT Report BD545 ) Figure 2 7. Excavated 0.406 m square x 6.1 m long jetted and grouted pile Figure 2 8 Three zones around an expanding cavity 3.25 Linear elastic zone Nonlinear elastic zone Plastic zone Cavity
46 Figure 2 9. Soil pile interaction S oil Z r m z r h Pile/shaft 0 r 1 2
47 CHAPTER 3 INDIVIDUAL RESPONSE OF JETTED AND GROUTED PILES This C hapter presents the analys is of previous experimental data on jetted and grouted piles, numerical modeling of grouting and axial load testing of the piles, and the proposed design methodology based on the findings of experimental and numerical analysis. A discussion of each follows 1 3.1 Analysis of the Previous Experimental Data on Jetted and Grouted Piles Detailed analysis of the results from the previous testing of Jetted and grouted piles (McVay et al. 2009) was performed. The test data of the following three piles were considered i n the analysis: 1) 0.153 m (6 in ) square x 2.44 m (8 ft ) long, 2) 0.203 m (8 in ) square x 2.44 m (8 ft ) long, 3) 0. 406 m (16 in ) square x 6.1 m long piles The final diameter of the side grout bulbs was 0.38 m (15 in) for the 0.153 m pile, 0.51 m (20 in ) f or the 0.203 m pile, and 0.914 m for the 0.406 m pile. 3.1.1 Test Chamber and Instrumentation m diameter x 10.67 m deep Lab was used for th e study (Figure 3 1) The benefits of a rigid wall chamber include: 1) control of the water table during pile/ shaft construction and testing, 2) replication of soil conditions for repetitive testing, 3) the use of instrumentation (vertic al and horizontal stress gages) in close vicinity to the piles/ shafts, and 4) the opportunity for soil excavation to expose grout zones 1 Major contents of Section 3.1, 3.2, and 3.3 are from the article: Thiyyakkandi, S., McVay, M., Bloomquist, D., Lai, P. (2012 Geotechnical and Geoenvironmental Engineering, ASCE. doi: 10.1061/(ASCE)GT.1943 5606.0000860. With permission from ASCE
48 For water level control, two 10 cm diameter slotted PVC pipes wrapped in filter fabric were located along the chamber wall to add or remove water from the chamber. Two 1.22 m (4 ft ) diameter and 13.72 m (45 ft ) long drilled shafts aligned with the centerline of the test chamber were used to provide reactions during the axial top down testing of the grouted pile s. Lateral and vertical soil stresses near the pile was measured during the installation, grouting and subsequent load testing using the earth pressure cells (Figure 3 2) installe d vertically and horizontally with in the test chamber 3.1.2 Test Soil Properties and Test Ch amber Soil Preparation The soil used in the test chamber was typical Florida silty sand (A 2 4), with a grain size distribution given in Fig ure 3 3 The fines in the soil were classified as non plastic. Minimum and maximum dry densities of the silty sand w ere 14.5 kN/m 3 and 18.1 kN/m 3 (92.2 and 115.2 lbs/ft 3 ), respectively. Direct shear tests performed on the soil at minimum and maximum dry densities produced peak angles of internal friction of 31 o and 36 o respectively. Figure 3 4 presents the shear force vs. shear displacement curves from the direct shear tests at maximum dry density. The large strain/critical state friction angle ( c ) of the soil was 31 o The soil was placed in the chamber in 0.6 m lifts. In each lift, the soil was allowed to free fall into the test chamber and was leveled to produce an un compacted layer. Each lift was then compacted with a vibratory plate compact or for 2 3 minutes, starting from the chamber boundary towards the center in a circular motion pattern. The moisture content of the soil was in the range of 5 7%. During soil placement, stress gages were placed at various depths and radial distances from t he expected pile locations to measure the radial and vertical stress changes within the soil. Density
49 measurements using the core cutter method/nuclear density gage on all the lifts revealed a mean dry density of 15.99 kN/m 3 with a coefficient of variatio n (COV) of 2.66% Based on the mean dry density, the relative density of compacted soil was about 47.5%. While filling the test chamber, a number of hand cone penetrometer tests were also performed on each compacted lift. For all the lifts, cone tip resist ances varied from 2450 2941 kPa (25 30 tsf) at 0.3 m (1 ft) depth to 4903 6864 kPa (50 70 tsf) at a depth of 1.2 m (4 ft). Using the cone tip resistance values, an average relative density of 56% was obtained using the expression (Equation 3 1) suggested b y Jamiolkowski et al. (2001). (3 1) where D r = relative density (%) q t = cone tip resistance atm = atmospheric pressure (1 bar = 100 kPa) vo = effective overburden pres sure Pressuremeter testing (PMT) was conducted in the test chamber after soil preparation to estimate the cylindrical cavity limit pressure in the soil at two different depths, 0.91 m (3 ft) and 1.83 m (6 ft), which corresponded to the middle of the top s ide grout bag and bottom side grout bag respectively for 2.44 m (8 ) long piles Figure 3 4 displays the expansion pressure vs. volume curves from the pressuremeter tests in the test chamber. The average limit pressure determined at depths of 0.91 m and 1. 83 m were 415 kPa and 520 kPa, respectively.
50 3.1.3 Residual Horizontal Stress Around Pile As mentioned earlier, the soil stresses within the test chamber were monitored during pile jetting, grouting and axial top down test Soil stress measurement before and af ter pile jetting revealed that considerable decrease in lateral soil stress occurred close to pile (Table 3 1 ) due jetting which is attributed to the soil disturbance around pile caused by jetting process. At larger radial distance, lateral soil stress ch ange was negligible. The measured grout pressure and grout volume during side and tip grouting of the piles are summarized in Table 3 2. The soil stress measurements during grouting showed that the radial and vertical stress around the pile increased durin g grouting and quickly diminished to a stable value after pumping ceased. The stress decrease after pumping was attributed to elastic unloading occurring around the pile and the incompressible nature of the grout. Evident from the measured residual lateral soil stress as shown in Table 3 1 a significant increase in stress occurred near the pile which decreased rapidly away from the pile Figure 3 5 presents the measured residual horizontal stresses around the final grouted piles over time. From the negligi ble changes in stresses with time, it was expected that the axial capacity of the pile would not change with time. However, more studies are required to quantify the possibility of stress redistribution (e.g., one year) due to creep, aging, etc. (Bullock e t al. 2005). Numerical modeling of the experimental study was performed to obtain more information on the soil stress es and displacements history around the piles during grouting and top down testing The measured grout pressures, grout volume, and residua l soil stresses after grouting were used to control the numerical modeling of the pile which is discussed in the S ection 3.2
51 3.2 Numerical Modeling of Jetted and Grouted Piles A two dimensional finite element program, PLAXIS 2D was used to model the install ation and top down loading of the pile. The pile and soil in the test chamber were simulated with an axisymmetric model using 15 node triangular elements, as shown in Figure 3 6 Along the chamber wall, both the radial and vertical displacements were restr icted. The bottom boundary was placed sufficiently far away from the pile such that its influence was negligible. The actual square pre cast pile was modeled as an equivalent circular pile with the same cross sectional area. 3.2.1 Material Models The sand was m odeled with the Hardening Soil (HS) constitutive model (described by Schanz et al. 1999, coded in PLAXIS), and the pile was modeled as a linear elastic material. The HS model is an advanced hyperbolic model formulated in elasto plastic framework (Schanz et al. 1999). Schanz et al. (1999) verified the model for large deformation cavity expansion problem by simulating p ressuremeter testing in loose sands. The model uses the Mohr Coulomb strength in the case of shear yield surface, which is governed by frictio n angle, cohesion, c and dilation angle, but has an additional yield cap to model the irreversible plastic strain due to isotropic loading The model uses three different Moduli : E 50 from primary deviatoric loading, E ur from elastic unloading/reloa ding and E oed from primary oedometer loading, to describe soil stiffness T he model also accounts for the stress level dependency of stiffness parameters according to a power law which is controlled by a dependency parameter m The material parameters use d for the sand and pile in the analysis are given in Table 3 3 E 50 ref was determined as secant modulus corresponds to 50% mobilization of shear strength from the standard drained tr ia xial deviatoric stress axial strain curve (Schanz
52 et al. 1999). PLAXIS default settings of E oed ref = E 50 ref and E ur ref = 3 E 50 ref were used in the analysis. Stress level dependency parameter, m was taken as 0.5 based on the recommendation of Anderson and Townsend (2005) and the PLAXIS authors. Anderson and Townsend (2005) r eported that the variation of m has little effect by investigating the HS model parameters for Florida sand. Elastic modulus of the concrete (pile) used in the analysis was evaluated as for c = 27.6 MPa (MacGregor 1992). Although t he cohesion of the silty sand used in the study was zero, a small cohesion value (0.345 kPa or 0.05 psi) was used in the analysis for the numerical stability of the model, as suggested in the PLAXIS manual. 3.2.2 Simulation of Grouting and Top Down Load Test Th e analysis began with the precast pile embedded at the installation depth ( i e ., the jetting process wa s not simulated) as PLAXIS 2D currently does not allow simulation of the jetting process. The grout zone and membranes were initially characterized as a soft linearly elastic zone ( E = 689.5 kPa or 100 psi and = 0.25), as shown in Fig ure 3 6 The zones were locat ed at the top and bottom of the pile as well as at the pile tip ; these zones represent loose soil occupying the wea k zone formed around the sh aft and tip during the jetting process. The grouting process was simulated by applying positive incremental volumetric strains to these elastic zones in steps which is the PLAXIS recommended approach to the simulation of grouting [ Ni and Cheng 2010] The sequence of the simulation was the same as that of the actual grouting of the pile (i.e., first side grouting and then tip grouting). The expansion of each elastic zone was controlled by the measured final grout volume (Table 3 2 ). In the PLAXIS analysis, the elastic unloading immediately after grouting was simulated by applying volumetric
53 contraction to the respective elastic zones a fter the application of the volumetric expansion representing grouting. The amount of volumetric contraction was controlled b y the magnitude of the residual stresses around pile, (i.e., residual horizontal stresses measured in the chamber tests near the membranes). Subsequently, each elastic zone was replaced by the linearly elastic concrete material to represent the hardened gr out. Because the grouting process is a large strain problem, the Updated Mesh Option was used in the analysis. The Updated Mesh analysis in PLAXIS is a calculation procedure based on the Updated Lagrangian formulation (Bathe 1982) to account for large defo rmation influences. The program updates the finite element mesh and stiffness matrix as the calculation proceeds to reassess stresses and strains within the mesh. Fig ure 3 7 shows a typical deformed mesh after the simulation of both side and tip grouting. The maximum expansion pressures for the top zone, bottom zone and tip of the pile s from PLAXIS analysis are given in Table 3 4 The grout pressures measured in the experimental study (Table 3 4 ) were slightly higher than the respective expansion pressures determined from PLAXIS. The difference was attributed to the resistance of the semi rigid membrane confining the grout zones. It should be noted that no interface elements were used at the boundary between the sand and membrane, as the semi rigid membrane around the grout bulb was extremely rough and the soil did not debond or slide relative to the membrane. The direct shear testing of membrane soil interface at test chamber soil density (16 kN/m 3 ) gave an interface friction angle of 29 o which is close to the friction angle of soil (31 o ). Therefore, in the FEM model, the interface friction angle was set equal to that of the sand (i.e., no soil strength reduction).
54 After simulation of the grouting process, the analysis modeled the pile load test by activati ng incremental distributed loads on top of the piles. The load displacement response of the piles obtained from numerical analysis was quite comparable with the measured response, as shown in Fig ure s 3 1 2 3 1 3 and 3 1 4 3.2.3 Lateral S tress D istribution and L a teral S oil D isplacement D uring P ile G routing One major concern with all grouting is the potential undesirable effects on adjacent structures due to soil movement and increased radial stresses generated by the grouting process. In geotechnical engineering, the process of grouting a pile in situ falls within the study of cavity expansion theory. Research studies by Carter et al. (1986), Yu and Houlsby (1991), Yu (2000), Salgado and Randolph (2001) and Salgado and Prezzi (2007) have contributed significantly t o the understanding of the cavity limit pressure and principal soil stresses in the vicinity of an expanding cavity. Grouting along the pile resembles cylindrical cavity expansion, and the tip grout bulb takes the shape of a spherical cavity. The radial st ress distribution within the plastic zone around an expanding cavity is inversely proportional to a power (for cylindrical cavity, power = (N 1)/N and for spherical cavity power= 2(N 1)/N where N is the flow number or passive earth pressure coefficient ) of the radial distance (r) from the center of the cavity (Carter et al. 1986; Yu and Houlsby 1991; Salgado and Randolph 2001). The radial stress is largest close to the cavity and diminishes with radial distance from the center of the cavity. The rate of d ecrease of the radial stress is high near the cavity, which is even more significant in the case of a spherical cavity (Yu and Houlsby 1991). The radial stresses measured around the spherical grout bulb during the tip grouting of 0.203 m square pile are pr esented along with stresses determined by FEM analysis and the
55 stress distribution using the analytical solution suggested by Yu and Houlsby (1991) in Fig ure 3 8 The variable along the horizontal axis is the ratio between the radial distance (r) and the cavity radius (a = 0.254 m). It can be seen that the measured stresses match reasonably well with the FEM and the analytical stress distribution. It should be noted that the radial stress diminished to 20% of the cavity stress at r/a =3 and 10% of the cav ity stress at r/a =5 (2.5 x bulb diameter). Similarly, the analytical solution for the lateral displacement around an expanding cylindrical cavity by Chai et al. (2005) reveals that the lateral displacement of soil in the plastic zone decreases hyperbolica lly with the radial distance. Shown in Fig ure 3 9 is the radial displacement observed during the numerical simulation of side grouting (lower zone) and tip grouting of a 0.203 m square pile in PLAXIS. It is evident from the Figure 3 9 that the displacement diminishes at a faster rate near cavity with radial distance, and the rate is even more striking in the case of tip grouting (spherical cavity expansion). The displacement was reduced to a considerably small value (less than 5 mm) at a radial distance of about 1.3 m (2.5x bulb diameter). Because the lateral stresses and displacements during grouting diminish rapidly with radial distance within a rather narrow zone close to the pile, any underground structures outside the zone are less likely to be affected by post grouting of the Jetted and grouted pile. In the present case the radius of the zone was about 2.5B (B = bulb diameter). The chamber boundary was 3.6B (1.83 m) in the 0.203 m square pile and 4.8B in the 0.152 m square pile. Moreover, no soil displ acement was observed near the chamber wall during the top down testing of the piles. Therefore, the results of the top down tests should be representative of field tests. However in the case of 0.406 m
56 square pile, the chamber boundary was 2B away only and consequently the chamber boundary may have influence d the test results 3.3 Design Methodology for Jetted and Grouted Piles Based on the results of the experimental study conducted by McVay et al. (2009) and the numerical analys i s a design methodology was p roposed for the Jetted and grouted piles. The methodology specifically includes approaches to estimate (1) the expected grout pressures during side and tip grouting of piles (2) the unit skin friction of the pile and (3) the load displacement response un der compression loading which are discussed below 3.3.1 Estimation of Jetted and Grouted Pile Grouting Pressures As mentioned earlier, side grouting resembles cylindrical cavity expansion, and the tip grouting is analogous to spherical cavity expansion. Cavit y expansion theory was subsequently used to predict the expected grout pressure during the installation of a jetted and grouted pile. In cavity expansion, t he radial stress at the wall of the cavity becomes the principal stress and is generally referred to as the limit pressure, as suggested by Menard (1957), who developed the p ressuremeter t est T he limit pressure may be used to predict the pump pressures required shaft and tip. In this work, the elastic perfectly plastic closed form solu tions of Yu and Houlsby (1991) and limit pressure charts provided by Salgado and Randolph (2001) were used to predict the expected grout pressure Another alternative to predict the expected grout pressure is to perform pressuremeter tests to obtain limit pressures at various depths at the site prior to the design of the jetted and grouted pile Table 3 4 shows a comparison of the measured pump pressures during grouting, expansion pressures from the FEM analysis and the limit pressures estimated using Yu an d form solution
57 (for a Mohr Coulomb material with a critical state friction angle, c = 31 o dilation angle, = 0 o = 0.25 ), Salgado and Randolph c = 31 o Relative density, D r = 50%) and p ressuremeter testing. It is evident from the Table 3 4 that the measured side grout pressures are slightly larger than the limit pressures at their respective depths. The increase may be attributed to the resistance offered by the semi rigid membrane and t he actual shape of the grout bulb (between that of a cylinder and a sphere) However the difference (20%) provides a reasonable prediction for the expected side grout pressure s required for installation It can also be seen from Table 3 4 that the measur ed tip grout pressure is less than the spherical cavity limit pressure at the specified depth. This difference may be due to the smaller tip resistance required to mobilize the side resistance of the pile. It is expected, for longer jetted and grouted pile s, that the tip grout pressure may be equal to the spherical cavity limit pressure, owing to the higher side resistance due to side grouting a longer (deeper) pile. 3.3.2 Estimation of Unit Skin Friction Besides estimating the expected grout pressures during in stallation, a designer must assess the expected axial capacity of the pile. Of major interest is the expected skin friction, which is verified during the construction process ( tip grouting ). However, if the tip grout pressure does not mobilize the full ski n friction of the pile, it is common to assume that the pile capacity is at least twice the measured grout tip load (Mullins et al. 2006). To estimate the expected side resistance, the normal stress (i.e., radial, hoop, and vertical) adjacent to the pile must be known. Yu and Houlsby (1991) estimated the
58 changes in the normal stresses around an expanding cylindrical cavity and reported that the normal stresses at the cavity wall increases 3 to 9 times above the in situ stress. During steady state cylindric al cavity expansion, radial stress ( r ) is the major principal stress, and the circumferential or hoop stress ( ) is the minor principal stress, which is close in value to the intermediate or vertical ( z principa l stresses during steady state cavity expansion touches the Mohr Coulomb c However, as mentioned earlier, the stress state around the pile changes due to elastic unloading immediat ely after grouting. It was observed from the experimental study and FEM analysis that radial and vertical stresses decrease and hoop stresses increase during elastic unloading after grouting. As a result, the radial stress/hoop stress becomes the major pri ncipal stress (depending on the extent of unloading), and the vertical stress be smaller (lower) than the strength envelope. However, for axial loading, the horizont al (radial) stress will diminish as the principal stresses rotate until the failure plane occurs vertically and the failure stress state occurs at the pole (Fig ure 3 1 0 ). Due to large stress changes, it is not possible to predict the magnitude of the horiz ontal normal stress at failure. However, it was found from the FEM analysis that the magnitude of the minor principal stress does not vary significantly with the mobilization of side resistance during axial loading; moreover, the minor principal stress at failure was nearly equivalent to the residual vertical stress ( vg ) after grouting. Figure 3 1 0 shows circle at failure, the pole under vertical loading, failure skin friction ( f s ) and the estimated vertical stress, vg
59 Using given in Figure 3 1 0 the value of the failure unit s kin fric tion ( f s ) was found in t erms of the critical friction angle ( c ) and effective vertical stress after grouting, vg (minor principal stress), (3 2) From the experimental and F EM analyses, the vertical effective stress ( vg ) at the grout soil interface of the pile was determined to be a function of depth ( h ), buoyant weight ( ) and the grout vertical effective stress coefficient, K g : (3 3) The proposed grout vertical stress coefficients, K g for sands are shown in Fig ure 3 1 1 The solid points were obtained from experimental testing and the open valu es from numerical analysis. Experimental K g hs of 1.8 m and 2.36 m in Figure 3 1 1 were obtained from the pressure cell data for the 0.153 m square pile. The vertical stresses before and after were measured using the earth pressure cells, which were about 0.152 m (6 in ) away from the pile. The pressu re cell at a depth of 1.8 m reported an initial vertical stress of 30 kPa and a vertical stress of 58 kPa after grouting, corresponding to a K g value of 1.93. Similarly, the vertical stress at 2.36 m was increased from 40 kPa to 68 kPa by grouting (i.e., K g = 1.7). The K g values: 1.98 and 2.25) at a depth of 1.37 m in Figure 3 11 were back calculated from the average unit skin friction of the 0.153 m square pile (42.8 kPa) and 0.203 m square pile (48.6 kPa) using Eq uations 3 2 and 3 3 Similarly, K g (1.6) at a depth of 3.1 m was back calculated from the average unit skin friction (75 kPa) of the 0.406 x 0.406 x 6.1 m jetted and grouted pile (Lai et al. 2010). It was assumed that the average unit skin friction develops at the middle o f the piles; hence, the back calculated
60 K g values represent that depth. The K g values obtained from numerical analysis agreed reasonably well with the experimental K g values ( c = 31 o ) and, consequently, for other friction angles ( c =33 o and 35 o ) ; the K g curves in Figure 3 11 were obtained from the numerical analysis results. Table 3 5 provides the prediction of the average ultimate unit skin friction ( f s ) and side resistance ( Q s ) of the piles u sing Equations 3 2 and 3 3, and Figure 3 11 Note that the K g values in the Table 3 5 were taken from the K g trend line ( c = 31 o ) in Figure 3 11 (not the back calculated values). 3.3.3 Load Displacement Curve for Jetted and G routed Pile s The pile soil load transfer approach proposed by McVay et al. (1989) was also used to estimate the load displacement relationship for the jette d and grouted precast pile. The nonlinear side shear at any point on the pile was assessed as a function of vertical deformation through the T Z curve ( Eq uation 3 4 ) Similarly the mobilized tip resistance as a function of tip displacement is given by the Q Z curve ( Eq uation 3 5 ) The load transfer functions are given as, T Z curve : (3 4) Q Z curve : (3 5) w her e r 0 = radius of pile r m = radius of influence zone = ) (Randol ph and Wroth 1978)
61 L = length of pile = r atio of G at L/2 to G at tip 0 = shear stress on pile soil interface, G = reloading shear modulus R f = ratio of failure shear stress to its ultimate R 0 = radius of tip bulb Q = mobilized tip load and R t = ratio of failure to ultimate tip resistance Q f = ultimate tip load Both Eq uations 3 4 and 3 5 are hyperbolic load transfer functions that depend nonlinearly on the shear modulus, G The ultimate skin fri ction, max used in ( Eq. 5) of the T Z curve is obtained from Equation 3 2 The ultimate tip load, Q f is obtained by multiplying the ultimate unit end bearing ( q b ) with the grout bulb tip area. The c orrelation between the spherical cavity limit pressure ( P L ) and ultimate end bearing pressure (Randolph et al. 1994) is used to estimate the unit tip resistance q b of the pile. ( 3 6) Comparison of l oad d isplacement c urve s : Fig ure 3 1 2 compares the predicted load displacement curve for the 0.152 m square pile using Eq uations 3 4 and 3 5 with the load test results and finite element analysis Similarly, Fig ures 3 1 3 and 3 1 4 show the 0.203 m square and 0.406 m square jetted and grouted piles, respectively. I t can be seen that the predicted load displacement response obtained using Eq uations 3 4 and
62 3 5 are quite comparable with both the measured response and results of the finite element analysis It is also evident from the shape of the load displacement c urves that the mobilized tip resistance occurs over a larger range of vertical displacements (2 mm to 30 mm). As identified by Lai et al. (2010), the reaction system for the 0.406 m pile load test was capable of developing only 1400 kN of compression axial load. 3.4 Full Scale Field Installation and Testing of Single Jetted and Grouted Piles F ull scale field construction and testing of two 0.71m square x 5.5 m long (28 in square x18ft long) Jetted and grouted piles were performed to validate the construction as well as the design estimates of torsion and axial resistance of the pile in typical Florida soils. S tatic top down test was performed on one of the piles and combined torsional and lateral load test on the other one For axial response comparison, a simil ar sized drilled shaft w as also constructed and tested in the same soil condition. The tests were performed in connection with the FD OT UF ongoing research project: BDK 75 977 41. The details of the test site soil investigation, the construction of j etted and g routed p iles and drilled shafts, the load tests and the test results are presented in the following sections. 3.4.1 Soil Investigation at t he T est S ite The site considered for the study FL Figure 3 15 shows the layout of test piles and drilled shaft along with reaction drilled shafts. A detailed subsurface exploration at the test site was performed by State Material Office, Gainesville Both in situ testing (SPT, CPT, PMT, and DMT) and laboratory soil testing (cl assification tests and direct shear tests) were conducted for the site. Standard Penetration tests were performed within the footprint of the test piles and shafts to a depth of about 3 times the width /diameter beneath the pile/shaft tips to assist
63 with th e design/construction. All the SPT borings revealed very high N values (exceeds 50) at a depth of 9 m through 14 m representing the existence of a hard stratum The water table at the test site was 2.9 m (9.5 ft) below the ground surface at the time of so il exploration, which seems to be perched on the hard stratum located at 9 m depth. Laboratory classification tests on the soil samples collected during SPT boring showed that, in general, the upper 0.91 1.22 m ( 3 4 ft ) consisted of clayey sand (SC) and l ow compressible clay (CL). SPT N values in this layer ranged from 3 to 8 The upper layer is underl ain by poorly graded fine sand with silt (SP SM) up to a depth of about 9 m SPT blow count in this sand layer varied from 3 to 34. From depths of 9 to 15 m very dense sand stratum with N value ranged from 51 100 exists. The hard sand stratum was followed by medium dense fine sand (N value: 17 to 33), which extended to the end of boring ( 21 m ). Figure 3 16 displays the SPT blow count (N) profiles and the Unif ied Soil Classification (USC) at the foot print of the test piles and drilled shafts along with their schematic drawings Shown in F igure 3 17 is the typical grain size distribution s for the silty sand at the site. Moisture content of the soils above water table ( 2.9 m ) varied from 1.5 to 20% with depth, whereas the soils below water table had reasonably uniform moisture content (25 30%) irrespective of the depth. Direct shear tests were conducted on the samples collected from all the bore holes at depths o f about 2.6 m and 4.6 m ( approximately corresponds to the center of grout bags of jet ted and grouted pile s ). The angle of internal friction varied from 27 o to 29 o Pressuremeter tests were performed in the footprints of the two jetted grouted piles at dep ths of 2.6 m and 4.9 m and the pressure volume curves are given in Figure 3 18. The p ressur emeter limit pressures (maximum value) at corresponding depth will
64 be the expected side grout pressure for each bag of jet ted and grouted piles as discussed in the d esign methodology for the pile ( S ection 3.3.1) 3.4.2 Design and Construction of Precast Piles U sed for Jetted and Grouted Piles 188.8.131.52 Structural d esign The precast pile used for jetted and grouted pile was 0.71m square in cross section. Steel reinforcement for the s ection was determined in accordance with ACI 318 08 C oncrete compressive strength of 34483 kPa ( 500 0 psi ) was used in the design Traverse steel reinforcement consist ed of #5 bars @ 5.08 cm (2 in ) spacing and longitudinal reinforcement wa s comprised of 16 nos. of #9 bars The pile section had a t orsional capacity of 1005 kN m flexural capacity = 1062 kN m, shear capacity = 1294 kN, and an axial capacity of 12428 kN. 184.108.40.206 Design and f abrication of g rout d elivery and j etting s ystems Figure 3 19 shows the schema tic diagram of jetted and grouted pile with the location of side grout bags and concrete cap Note that the use of concrete cap was to transfe r the load during the combined torsion and lateral test of the piles The design and construction of the concrete cap is discussed later. There were two 2.13 m (7 ft) long side grout bags with se parate grout delivery systems. In conformity with the previous studies ( McVay et al. 2009, Thiyyakkandi et al. 2012 ), 2.5 cm PVC pipes were selected for the grout delivery sys tems shown in Figure 3 2 0 Each of the grout pipes (entry & exit) had 7 pairs of holes with 7.5 cm center to center spacing ( F igure 3 2 0 ), evenly distributed (12.5 cm intervals) along the bottom 2/3rd of each grout bag 2.5 cm diameter gum rubber (6 mm thi ck) membrane cover ed each pair of holes
65 Figure 3 2 0 also shows the PVC piping fabricated for jetting of precast pile into ground. The jetting pipes was also be used for tip grouting. Diameter of the central jetting pipe was determined based on the guidel ines recommended by Tsinker (1988) with the flow rate equation (Equation 2 1) for sandy soil suggested by Shestopal (1959). Table 3 6 presents the estimation of diameter of the central (main) jetting pipe. Using Equation 2 1 ), require d flow rate for the present case was found to be 82.3 m 3 /hr. Based on the recommendations of Tsinker (1988) and Gabr et al. (2004) a flow velocity of 5 m/s was considered in the present case and required jet pipe diameter was estimated to be 7.5 cm (Table 3 6 ). The central jet pipe branched off into five 5 cm diameter pipes at bottom of pile for uniform distribution of water at tip. 220.127.116.11 Construction of p recast p iles and p reparation for j etting Figure 3 2 1 shows the reinforcement cage along with grout delivery systems, jetting system, and instrumentation Also shown in the Figure 3 2 1 ( middle ) are 4 bolts/ side attached via 0.95 cm thick x 6.35 cm wide steel plates to reinforcement cage. These threaded bolts we re used for attaching the side grout membranes to t he pile prior to jetting. The piles were instrumented with 5 concrete embedment strain gages as shown in Figure s 3 19 and 3 2 1 for monitoring the load distribution along the piles for the top down load test Figure 3 2 2 is the photograph showing the concre te placement of one of the piles. Average 28 th day compressive strength of the poured concrete was 55634 kPa. After sufficient curing, the formwork was detached and excess concrete at the location of side grout exit (covered with gum rubber) was removed T hen the precast piles were prepared for jetting, which include: (1) flushing of each grout delivery systems to ensure proper function, (2) t esting the membrane, (3) a ttaching top and
66 bottom side grout membranes, (4) a ttaching nozzles for jetting/tip grouti ng system and (5) t esting of nozzles to ensure uniform water distribution at tip E ach grout delivery systems (both top and bottom) was connected to the city water supply and flushed to ensure that the systems were working properly (Figure 3 2 3 ) A pressu re of about 48 kPa ( 7 psi ) was needed for water to exit through the ports after expanding the gum rubber cover The membrane used for the side grout bags was 45 mil thick reinforced polypropylene geomembrane (RPP). The membrane was first tested for interf ace soil a dry density of 16 kN/m3 (102 pcf) and a moisture content of 7.5%. The soil used for the tests was typical Florid a silty sand o Tensile strength of the membrane was estimated by conducting tension tests on a number of 3.18 cm ( ) wide strips and the average tensile strength of the membrane was 22138 kPa. T o form the grout bags, the membranes were first rolled to a cylinder of 1.22 m diameter (design diameter of grout bulb ) with an overlap of 15 cm and then the overlaps were heat seamed. For each grout zone, the total length of the required membrane wa s the su m of the length of side grout zone (2.13 m) additional length for expansion outward during grouting, and width of steel plates for attaching the membrane to pile at both ends (i.e., 2.13 + 0.508 +2x 0.075 = 2.788 m + 0.075 m extra = 2.863 m ). Each vertically sea med membrane was cut to a length of 2.863 m A grid of small holes ( 1.5 2 mm ) at 15 cm spacing were drilled into the membrane, which is expected to allow the grout seepage through the membrane during grouting process and thus develop
67 adequate bonding bet ween grout bag and surrounding soil In order to attach the membrane to the pile, both vertical and horizontal pleats had to be placed in the membrane. Note that the vertical pleats will reduce the circumference and one horizontal pleat will reduce the hei ght of the membrane. For each pile, the horizontal pleat was first made in the middle by folding the membrane roll. Then, the membrane was positioned at the l ocation of grout zone Next, the membrane was attached to the pile in such a way that a vertical p leat of about 11 cm wide was made in the middle of each face. The membrane was secured to the pile by means of steel plates and threaded rods embedded in pile ( Figure 3 2 3 ). After the attachment of both top and bottom bags, t he nozzles were attached at th e exit of all the 5 cm jet pipes at the tip of the piles ( Figure 3 2 3 ). The nozzle pattern for all the outer pipes was same; four 1.25 cm diameter holes, as shown in Figure 3 2 5 F or the central pipe, the nozzle pattern consists of four 1.25 cm diamete r holes and one 1 cm diameter hole in the middle. The nozzles were tested to ensure unifor m water distribution at the bottom ( Figure 3 2 3 ). Test was performed by connecting the central jet pipe to the city water supply. The maximum water pressure recorded by pressure gage was less than 35 kPa As evident from the F igure 3 2 3 flow distribution was virtually uniform even under small pressure. 18.104.22.168 Jetting of p recast p iles After the preparation (i.e., attaching the membranes and nozzles) the piles were jetted i nto the ground Figure 3 2 4 shows the setup for the pile jetting. In order to reduce water loss (e.g., percolation) and minimize water requirement during jetting, the water used for the jetting was re circulate d. First of all, 2.1 m diameter x 1 m deep hol es were made and temporary surface casings were installed at the location of both piles to
68 collect the water coming up during the jetting process ( Figure 3 2 4 ) A 50 ton crane was used for positioning and holding the pile during the jetting process The pr essurized water for jetting was provided to the test pile from a water tanker through a 15 cm high pressure jet pump (max. flow rate = 6 m 3 /minute and max. pressure = 1269 kPa ), Figure 3 2 4 The pump was equipped with flow meter and pressure gage to monito r the flow rate and pressure respectively. Jetting initiated with flow of water from the water tank through the pump to the test pile with a flow rate of 1.5 m 3 /minute and a pressure of about 900 950 kPa. Subsequently the test pile was lowered with the cra ne as penetration occurred. The pile was allowed to penetrate with its own self weight by releasing the weight steadily from the crane A hydraulic trash pump (maximum flow rate = 4.9 m 3 /minute and pressure = 450 kPa) was used to pump the water collected i n the surface casing back to the tanker ( Figure 3 2 4 ) at a flow rate of nearly equal to the jetting flow rate ( 1.5 m 3 /minute ). Total water loss (percolation) during the jetting of two piles was a pproximately 3.75 m 3 After jetting of each pile, the casing was pulled out and soil was backfilled around the pile The n oise and ground surface vibration s generated during the pile jetting operation were recorded using Vibration and Overpressure Monitors (triaxial geophones and overpressure microphones). Analyses of noise and vibration data are presented later. 22.214.171.124 Design and c onstruction of c oncrete c ap for j et ted and g routed p iles A reinforced concrete cap w as required for transferring the forces and moments from the loading assembly (Mast arm structure) to the pil e during the combined torsion and lateral load testing. A precast concrete cap was chosen for one of the jet ted and grouted piles and a cast in place cap for the other Figure 3 2 5 depicts the schematic of
69 precast concrete cap pile connection. The starti ng point for the design of concrete cap was the dimension of inner square hole for accommodating the precast pile head. In case of precast cap, a 5 cm grout space/gap between the cap hole and pile was considered, which would provide adequate room for lower ing the standard grout tube that the industry use s From there, other dimensions and various components of cap (anchor bolts, flexural, torsional and shear reinforcement) were designed to meet various stan dard code requirements (ACI 318 08 AASHTO LRFD, AI SC 360 05, etc. ). The diameter and number of ASTM F1554 Grade 55 anchor bolts required to transfer the loads from Mast arm pole to foundation were determined by considering steel strength requirements of anchors in tension and shear (ACI 318 08, AASHTO L RFD, AISC 360 05) I t was found that 16 nos. of 3.75 cm ( 1.5 in ) diameter bolts we re sufficient in the present study. The optimum dimensions (outer width and depth) for the cap was selected considering the concrete breakout strength of anchor groups in sh ear, side face of anchors in tension, concrete pry out strength in shear, anchor bolt 1.74 m (60 in ) outer width and 0.77 m ( 30.375 in ) depth was found to be adequate f or transferring the forces and moments generated during the load tests without concrete breakout and side face blowout failure. A minimum clear cover of 7.5 cm was provided to all the reinforcements for all the exposed sides of cap in accordance with FDOT' s Structures Design Guidelines for LRFD ( FDOT 2002b ), considering the site condition as depth was determined by considering the development length for longitudinal reinforcement and spacing
70 be tween anchor bolts and longitudinal reinforcement in accordance with NCHRP (2003). The development length for # 5 rebars was determined to be 0.404 m using ACI 318 08 (2008) and the horizontal distance between anchor bolt and vertical reinforcement wa s abo ut 0.149 m ( 5.87 in ) Therefore the required embedded depth for anchors was estimated to be 0.628 m [0.404+0.149 + top clear cover ( 0.075 ) = 0.628 m ]. Steel reinforcement for concrete cap was calculated according to ACI 318 08 (2008 ). A concrete strength of 34483 kPa ( 5000 psi ) was used in the design Figure 3 2 6 displays the longitudinal and cross section of cap with reinforcement details. Note that the dimensions and various reinforcements for both precast and cast in place cap were the same. Only differen ce wa s the grout space/gap for the precast one. Estimation of grout tensile strength, shear strength and bond resistance at grout concrete interface of the precast cap pile connection using the strength properties of cementitious non shrink grout approved by FDOT (Qualified Products list) indicate d that the interface could safely transfer the forces from cap to pile during the load tests Figure 3 2 7 shows the placement and grouting (gap between the pile and cap) of precast cap The grout was filled from bo ttom to top (free flow) using a funnel and a hose as shown in Figure 3 2 7 Similarly, Figure 3 28 presents the c a sting of cast in place concrete cap 126.96.36.199 Side and t ip grouting of the p iles The side membranes of the piles were grouted only after allowing suffic ient time for the hydration of concrete cap installed at the top of piles (Figure 3 29 ) Theoretical grout volume required to fill a purely cylindrical side membrane was estimated to be about 1.325 m 3 ( 350 gal lon ) Since the membrane had vertical pleats on both ends to attach to the pile faces (reduced perimeter at both ends), the actual volume needed to fill the membrane will be less than the theoretical volume. A grout volume of
71 a pproximately 1.14 m 3 ( 300 gal lon) may be sufficient for each membrane. The g rout volume and sustained grout pressure were recorded throughout the grouting process for each membrane. The grout pressure was recorded at the pump and the top of the pile ( both inlet and outlet pressures ). Pile head and soil displacement were also monit ored throughout the grouting process (Figure 3 29 ) Figures 3 3 0 and 3 3 1 present the grout pressure (inlet pressure at pile head) versus the accumulated grout volume pumped in the top and bottom membranes of the pile 1 and 2 respectively. Table 3 7 compa res the measured grout pressures with the predicted grout pressures from the different approaches, such as Yu and Houlsby (1991) closed form solutions, Salgado and Randolph (2001) charts, as well as with the limit pressures from the Pressuremeter tests As evident from the Table 3 7 the measured side grout pressures were in good agreement with the predicted grout pressures After the hydration of side grouting, tip grouting of both piles were performed. The sustained grout pressure, volume of grout pu mped and vertical displacement of pile and surrounding soil were recorded throughout the grouting process. The tip grouting was controlled by the upward displacement of pile Specifically, the grouting stopped when the average upward pile head displacemen t exceeded 9.5 mm ( in) in combination with a steady or dropping tip grout pressure. Generally, 9.5 mm to 12.5 mm ( to in) of displacement is considered sufficient to fully mobilize the skin resistance on a pile in the literature. Shown in Figure 3 3 2 is the measured grout pressure versus volume of grout pumped during the tip grouting of both piles. Figure 3 3 3 displays the grout pressure versus pile head displacement plots In the case of Pile 2, as shown in Figures 3 3 2 and
72 3 3 3 there was an approxi mate linear increase of grout pressure with volume until pile began to move upward (2690 kPa), and grout pressure dropped off with further vertical movement (i.e. full mobilization of skin friction) In the case of pile 1, the grout pressure and volume inc reased until 1552 kPa, upon which pressure dropped, but grout volume below the tip still increased (Figure 3 3 2 ) At grout volume of 0.3 25 m 3 (Figure 3 3 3 ), the grout pressure started to increase and at grout volume of 0.53 m 3 upward movement had reached 9.5 mm (3/8 in) and grout pressure dropped suggesting that the full mobilization of skin friction had occurred. Table 3 8 presents a comparison of the measured maximum sustained tip grout pressures versus the spherical cavity expansion limit pressures at pile tip predicted using Yu and Houlsby (1991) closed form solutions, and Salgado and Randolph (2001) charts. It is evident from the Table 3 8 that the measured grout pressures were close to the predicted spherical cavity limit pressure at that depth, indicating that the maximum possible grout pressure (i.e., steady state expansion) had developed during the tip grouting of both piles. Specifically, the side grouting of the piles significantly improved the side resistance, i.e. the lateral stresses adjac ent to the piles upon which tip grouting leads to spherical cavity expansion The noise and ground surface vibration were also monitored during the top membrane and tip grouting of both piles, which is discussed below 3.4.3 Measured N oise and V ibration D uring P ile J etting and Grouting As identified earlier noise and ground surface vibration was measured during the pile jetting and grouting operation Vibration and Overpressure Monitor ( Instantel with triaxial geophone (velocity transducer) and overpressure
73 microphone were used for the measurement (Figure 3 3 4 ) The monitors were located at different radial distances from the pile location 188.8.131.52 Measured noise Figure 3 3 5 sho ws the location of construction equipments and vibration and noise monitors with respect to the pile location during the jetting process. Note that the layout of noise and vibration monitors was same for both piles, but the location of the construction equ ipment was different. The construction equipments used to jet the piles were: 1) high pressure jet pump, 2) crane, and 3) water recirculation pump, as discussed earlier. Therefore the major sources of noise in the jetting operation were the sound/noise emi tted from the motors of these equipments, and not from the jetting process itself According to hourly equivalent steady state sound level (L eq ) should be limited to 67dBA in residential, hospital, school, picnic, an d recreational areas and 72dBA in developed lands (commercial and industrial measurement in decibel (dB) adjusted/weighted to match the sensitivity of human ear. Figure 3 3 6 displays the noise measured at various monitor locations during the jetting process T he noise in the vicinity of pile location was in the range of 85 92 dBA, which is attributed to the operating noise of water recirculation pump and high pressure jet pum p. It can be seen from the Figure 3 3 6 that t he noise levels were nearly the same for both piles even though the locations of equipments were different (Figure 3 3 5 ) This is due to the fact that resultant noise level at a location due to multiple noise s ources is dominated by the highest individual noise level (as sound level, dB or dBA is logarithmic ) It is evident from the Figure 3 3 6 that for the distance beyond 21 m, the
74 noise level is less than 72dBA (FHWA NAC criterion). T he noise can be further re duced by shielding the equipments (jet pump and water recirculation pump) and locating the pumps away from any sensitive building/s tructure under consideration. ( FHWA 2006 ) database suggests an A weighted maximum sound level ( L max ) of 95dBA at a distance of 15.2 m (50ft) for any impact pile driving operation. Since the noise levels decrease with the logarithm of distance from the source, the corresponding hourly equivalent sound level ( L eq ) at various distances can be ob tained using Equation 3 7 ( FHWA 2006 ). Figure 3 3 6 also displays the estimated variation of noise levels with distance for an impact pile driving operation (using Eq uation 3 7 L max = 95 dBA and U.F (%) = 100 ) In the case of impact/dynamic pile driving, i t can be found using Equation 3 7 and an L max value of 95 dBA that the noise level is less than 72 dBA (FHWA 2006 ) at a radial distance beyond 213 m (700 ft) Therefore it can be concluded based on limited data that the noise generated during a pile jettin g operation is much less than a dynamic pile driving as expected. (3 7) Where, D distance from source, U.F (%) time averaging equipment usage factor During the grouting of the piles, the grout pump/diesel generator was the only construction equipment u sed and hence the only source of noise. In all the cases (i.e., side and tip grouting) the grout pump was in the range of 8.5 15.2 m away from the pile and in the opposite side of noise and vibrati locations. Figure s 3 3 7 and 3 3 8 show the nois e measured at different locations during the top membrane and tip
75 grouting respectively It is evident from the Figures 3 3 7 and 3 3 8 that the noise residential area) and he nce the noise during the grouting process is not critical as long as the source (grout pump) is at least 16 m away from the location under consideration. 184.108.40.206 Measured ground surface vibration Also of significance is the induced ground vibration during the cons truction. Ground motion may cause structural and architectural damage to nearby structures Triaxial geophones (velocity transducers) were used to measure the three orthogonal components (transverse, vertical, and longitudinal) of particle motion at differ ent radial distance during the pile jetting and grouting operation. The resultant particle motion was determined as the vector sum of three orthogonal components. Shown in Figure 3 39 are the peak particle velocity profiles for pile 1 and 2 during the jett ing process The peak particle velocity was higher (0. 48 mm /s) for pile 1 and it was attributed to the proximity of jet pump to the pile as shown in Figure 3 3 5 However, beyond 9.1 m (30 ft), the particle velocity was negligible. For pile 2, the peak part icle velocities were similar (1.8 mm/s) at 3.66 m and 9.75 m due to the similar distance from the jet pump to the pile (see Figure 3 3 5 ) At larger distances, peak particle velocity was higher in the case of pile 2 than pile 1, which is attributed to the d ifference in location of the jet pump as identified from Figure 3 3 5 Table 3 9 presents the maximum vibration level recommended by AASHTO Designation R8 81 (AASHTO 198 2 ) to avoid structural damage. It is evident from Figure 3 39 that for distance greater than 6. 5 m peak particle velocities were less than the minimum limiting velocity ( 2.5 mm/s or 0.1 in/s) suggested by AASHTO Designation R8 81 for historical or critical structures.
76 Figure 3 4 0 presents the peak (maximum) particle velocity measured during the side and tip grouting of the piles It is clear from the Figure 3 4 0 that vibrations during the process were negligible; much less than minimum limiting velocity suggested by AASHTO Designation R8 81 ( Table 3 9 ). 3.4.4 Axial Top Down Testing of the Jetted and Grouted Pile After waiting approximately four weeks following the tip grouting of the piles, static top down testing of jet grouted pile 1 was performed in accordance with ASTM D 1143/D 1143M 07 Figure 3 41 shows the setup for the load test with all t he instrumentations. The reaction for the static load test was provided by two 1.22 m diameter x 12.2 m long reaction drilled shafts installed on either side of the test pile (5.5 m away) Two 12.2 m long beam girders (Figure 3 41) were used for transferri ng the load to the reaction drilled shaft during the load test An 8896 kN ( 2000 kips ) capacity hydraulic jack was used for applying the load ( F igure 3 41 ). The applied load was measured using a n 8896 kN load cell. The vertical displacement of the pile was monitored using the digital levels, digital dial gages, and mirrored scale (Figure 3 41 ). The digital dial gages were attached to a wooden reference beam as shown in Figure 3 41 The upward displacement of the reaction drilled shafts was also monitored th roughout the test. The data from the strain gages embedded within the test pile was acquired using the National Instruments data acquisition system. The data was later used to identify the load distribution along the pile and used to separate the skin and tip components. The load was applied in 111 kN (25 kips) increments and each increment was kept for a constant time interval of 10 minutes in conformity with ASTM D 1143. The loading could not be continued beyond 1557 kN ( 350 kips ) due to failure of one o f the
77 reaction drilled shafts (u pward movement, i.e. pullout). The maximum displacement observed on the top of the test pile was only 4 mm (0.15 in ). The applied load was then removed in five decrements. Shown in Figure 3 42 is the strain measured at diffe rent locations: above the top bag, between the bags, and below the bottom bags; during the load test. The load distribution alone the pile, and skin and tip contributions were estimated using the measured strains and applied top load. Figure 3 43 presents the total load, mobilized tip load, and mobilized skin resistance versus top displacements during the load test. Shown in Figure 3 44 is the load distribution along the pile du ring the incremental loading It can be seen from the F igures 3 43 and 3 44 that that the side resistance of the pile was not fully mobilized during the load test. The total mobilized side resistance of the pile was 1254 kN ( 282 kips ) The unit skin and ultimate side resistance of jet ted and grouted pile was estimated using the propos ed prediction methodology ( S ection 3 3 ) and using the tip grout pressure data Table 3 1 0 presents the prediction of the average ultimate unit skin friction ( f s ) and side resistance ( Q s ) of jetted and grouted pile using the proposed methodology. The averag e unit skin friction for each membrane was determined at the average depth using the K g plot ( F igure 3 11 ) and Equation 3 2 The side resistance was obtained by multiplying the average unit skin friction with the pile surface area. Note that the diameter o f the bulb was estimated by assuming a purely cylindrical shape with volume equal to the volume of grout pumped. Second approach was based on the field reported tip grout pressure (i.e., construction approach) The ultimate axial skin friction of a pile sh ould be equal to the
78 maximum tip grout pressure times the effective tip area. The effective area is assumed as the area of circle with diameter equal to the diagonal distance of precast 2 = 0.7945 m 2 ]. The skin resistance based on the tip grout pressure is found to be 1589 kN (2000 kPa x 0.7945 = 1589 kN; Note total capacity: skin+ tip 3178 kN), i.e. sufficient load has not been applied to the pile to mobilize fu ll skin friction or tip resistance Table 3 1 1 provides a comparison of the measured skin resistance with predicted skin resistance using different approaches. It is evident from the Table 3 11 that the measured skin resistance for the top membrane was in reasonable agreement with the predicted value based on K g approach However the proposed methodology under predicted the skin resis tance of the bottom membrane. This may be attributed to the actual shape of the bottom side bulb formed which is currently u nknown because the pile is not excavated It is also clear from the Table 3 1 1 that the measured tip grout pressure predicted the ultimate skin resistance quite well. 3.4.5 Construction and Testing of Comparison Drilled shaft 220.127.116.11 Construction of test drilled shaft The drilled shaft considered for the axial resistance comparison was a 1.22 m diameter x 5.5 m long shaft. The shaft was instrumented with 8 sister bar strain gages (2 gages diagonally opposite in 4 different levels: 1.68 m, 3.35 m, 4.42 m, and 5.03 m) Wet construction method was employed for the construction of the drilled shaft. Mineral slurry (bentonite clays) was used in the excavation to provide stability. A truck mounted drill rig was used for excavation. Figure 3 4 5 shows the lowering of reinforcing cage in to the shaft hole using a truck mounted hydraulic crane and the concrete placement The average 28th day compressive strength of the concrete determined from the testing of the sample cylinders was 48100 kPa.
79 18.104.22.168 Axial top down testing of drilled shaft Static top down load test on the drilled shaft was performed for comparison with jetted and grouted pile. The load test setup was the same as that for the jetted and grouted pile. As in the case of jetted and grouted pile, the reaction for the load test wa s provided by two 1.22 m diameter x 12.2 m long reaction drilled shafts. The vertical shaft displacement monitoring included the digital dial gages and mirrored scale with wire line reference. The load was applied in 89 kN (20 kips) increments with a time interval of 10 minutes and continued until a vertical displacement of approximately 12.5 mm (0.5 in) was observed at the shaft head. Then the load was removed in four approximately equal decrements. The water table during the load test was 3.05 m below the ground surface. The sister bar strain gages installed at the different elevations of the test drilled shaft were also monitored throughout the test. The measured strains at different elevations were then used to calculate the loads at the different elevat ions. Shown in Figure 3 4 6 is the load distribution along the shaft during the application of each load increments. It is clearly evident from the Figure 3 4 6 that the side resistance of the shaft was fully mobilized ( 1068 kN lines become parallel) prior to peak test load ( 1423 kN ). Figure 3 4 7 depicts the total load vs. top displacement of the shaft along with separated skin and tip contributions. The ultimate skin resistance of the drilled shaft was found to be 578 kN 3.4.6 Comparison of the Response of Jett ed and Grouted Piles with Drilled S haft S hown in Figure 3 4 8 is the total load displacement response of the jetted and grouted pile and similar sized drilled shaft. It is evident from the Figure 3 4 8 the axial resistance of the jet ted and grouted pile was much greater than that of the similar sized
80 drilled shaft Table 3 1 2 presents the mobilized un it skin frictions for each zone of jetted and grouted pile and a comparison with the maximum unit skin obtained during the drilled shaft test. It is clear from the Table 3 1 2 that unit skin friction for the jet ted and grouted pile is much greater than that of drilled shaft, especially in the bottom zone (2.6 times). The superior unit skin resistance for jetted and grouted pile is attributed to the larger horizont al stress close to the pile subsequent to side grouting (cylindrical cavity expansion). 3.4.7 Combined Torsion and L ateral Load Testing of the Jetted and Grouted Pile 22.214.171.124 Design and fabrication of M ast arm a ssembly Past torsion testing of jetted and grouted pile in a large test chamber environment ( McVay et al. 2009) showed that the pile possess very high torsional resistance and hence could be used as the foundation for Mast arm structures supporting highway signs and signals. Since the eccentric dead load of a sta ndard FDOT Mast arm assembly itself generate axial load and moment ( M x ; about the axis parallel to arm axis) on the top of foundation, a real Mast arm structure need to be used for the testing. One of the longest (23.77 m or 78 ft long) FDOT Mast a rm avail able is E7 type and corresponding pole is T6 type (FDOT Desi gn Standards Index No. 17743). FDOT Design Standards prescribe a 1.22 m diameter x 5.5 m long drilled shaft (4 ft x 18 ft) to support an E7 T6 Mast arm assembly. Table 3 13 shows the forces and mo ment s generated on foundation top for the E7 T6 Mast arm assembly subjected to a design wind speed of 130 mph (FDOT Mathcad 14 spreadsheet : MastArm v4.3 ). The orientation of the coordinate system considered is shown in Figure 3 49. It was found using the FDOT MastArm v4.3 that the standard E7 T6 assembly is not capable of carrying torque ex ceedin g 413 kN m (305 kip ft). Because the jetted and grouted pile
81 was expected to have a torque resistance more than 413 kN m, a new Mast arm structure had to be designed and fabricated for the testing of the pile to failure Accordingly a Mast arm assembly wit h the dimensions shown in Table 3 14 was designed sheet: MastArm v4.3 to carry a maximum torque of 75 9 kN m (560 kip ft). The spread sheet had t o be slightly modified to incorporate a point lateral load on arm instead of wind load. Note that the new Mast arm develops similar axial load and moment on the pile as an E7 T6 assembly due to its self weight (Table 3 13 ). Also the forces, moments at top of pile, due to 33 kN (7.4 kips) lateral load on the new Mast arm at an eccentricity (standoff distance) of 10.67 m (35 ft) are comparable with the load combinations for an E7 T6 type assembly at a design wind speed of 130 mph as a shown in Table 3 13 Fig ure 3 50 shows the Mast arm assembly (i.e., arm, pole, and connection bracket) fabricated in accordance with design. 126.96.36.199 Test s etup, i nstrumentation, and l oad t est p rocedure A crane with an axial capacity of 747 kN ( 75 ton ) was used to lift and set the Mast ar m assembly on the pile (Figure 3 5 1 ). After setting and orienting the Mast arm assembly properly, the bottom flange of the pole was bolted to the anchor rods embedded in the concrete cap, while the crane supported a significant portion of the dead weight o f the assembly (Figure 3 5 1 ). Three different types of instrumentations were used for the monitoring of rotation and translation of pile during the load test: (1) two Total Stations, (2) two sets of string pots (4 pots in each set), and (3) four digital di al gages. Figure 3 5 2 displays the locations of Total Stations and the targets on the Pile Mast arm assembly. As shown in Figure 3 5 2 the reflective tapes were used as the targets, which were glued to eye bolts attached to the end of 3.75 cm diameter stee l pipes projecting outward from the pole.
82 The targets were at the two different elevations ( 4 at each elevation; Figure 3 5 2 ); the first set was 0.15 m (0.5 ft) above the bottom flange and the second set (top) was 1.5 m (5 ft) above the bottom set Figure 3 5 3 shows the schematic arrangement and the actual placement of the string pots at the test sit e. The string pots sets (4 in each set) were also placed at the same elev ation of Total Station targets. The string pots with supporting frame were kept outside the influence zone of pile (i.e., 5 x diameters away from pile). National Instruments data acquisition system was used to record the string pot data. Shown in Figure 3 5 4 is the placement of digital dial gages at the elevation of bottom Total Station targ ets and string pot set (i.e. 0.15 m above bottom flange). The gages were supported by a wooden reference beam as shown in Figure 3 5 4 The depth of water table on the day of testing was 2.62 m (8.6 ft). The combined torsion and lateral loading of the pile was performed by applying lateral load on the mast arm at an eccentricity of 10.67 m (35 ft) using a crane (Figure 3 5 5 ). The location of crane was about 40 m awa y from the arm. A surveying level was used to ensure that the cable was horizontal during the loading. The crane applied the lateral load at an increment of 2.22 kN (0.5 kips) by pulling on the arm ( Figure 3 5 5 ). The applied lateral load was measured using an 89 kN (20 kips) capacity tension load cell attached to the cable (Figure 3 5 6 ). Each incr ement was kept for a uniform time interval of 5 minutes. The loading could not be continued beyond 54 kN (12 .17 kips) due (tag line) Subsequently, the load was removed in seven decrements with a time interval of 5 minutes. 188.8.131.52 Analysis of r esults Table 3 15 presents the forces and moments acting on the pile when the arm was subjected to the maximum lateral load of 54 kN As given in Table 3 15 the
83 foundation was under a combination loads and moments ( V x V y M z M x T ) during the load test Figure 3 5 7 displays the torque vs. rotation response measured during the combined to rsion and lateral load testing. It can be seen from the Figure 3 5 7 that the rotations measured using different instrumentations (Tot al Stations, string pots, and dial gages) were the same. The maximum value of torsional resistance mobilized during the load test was 577.6 kN m ( Table 3 15 and Figure 3 5 7 ) and corresponding measured rotation was only 1.45 o which indicates that the torsi onal resistance of the pile was not fully mobilized during the test. Shown in Figure 3 58 is the components of lateral displacement along the direction of lateral load (X axis) and axis of arm (Y axis). It is evident from the Figure 3 58 that the displacem ent component along the arm axis was the larger in magnitude up to a lateral load of 36kN and afterward the displacement along the load direction became the major component Figure 3 5 9 presents the lateral load vs. resultant lateral displacement response of the pile. The maximum lateral displacement observed was only 15 mm. Interestingly, there was significant elastic rotation and translation rebound during the unloading phase of the test as shown in Figures 3 57 and 3 5 9 Specifically more than 50% of the rotation and translation during the loading phase was elastic in nature It is also evident from the Figure 3 60 which shows the torsional cracks and gaps due to combined rotation and translation after maximum loading and full unloading. The gaps formed during the loading phase were almost recovered during the unloading phase. The torsional resistance mobilized during the test was subsequently compared with the values predicted using the different approaches (Table 3 16 ). The torsional resistance was est imated by multiplying the surface area of the pile with the unit skin
84 friction values predicted using the proposed methodology (Equation 3 2) and the tip grout pressure. The surface area of the pile was estimated by assuming that purely cylindrical shape d bulbs with volume equal to the volume of grout were formed around pile. Note that the torsional resistance contribution due to the pile tip is not considered in the prediction. It can be found from the Table 3 16 that the methodology presented in Section 3 .3 under predicted the torsional resistance. It may be attributed to the actual shape of the side grout bulb formed and the fact that tip contribution was not considered in the prediction. Since the torsional resistance of the pile was not fully developed during the test (rotation = 1.45 o ), it is expected that the ultimate torsional resistance (i.e., 8 10 0 rotation) of the pile will approach the value predicted using the tip grout pressure. The combined torsional and lateral load testing of the jetted and grouted pile in typical field scenario shows that the torsional resistance of the pile ( 577.6 kN m) even under very small rotation is much greater than the torque (351 kN m) generated on an E7 T6 Mast arm structure during a design wind speed of 130mph ( e. g., hurricane). Noteworthy that it is even much higher than the torsional capacity of E7 T6 Mast arm structure s (413 kN m). Thus the test suggests that the new pile is an appropriate foundation system for Mast arm structures supporting highway signals and signs considering both constructability and resistance in typical cohesionless soil conditions.
85 Table 3 1 Measured lateral soil stress change due to jetting and grouting Pile Initial stress gage location Lateral stress De pth (m) Radial distance from pile surface ( in terms of pile width, B) Before jetting (kPa) After jetting (kPa) % of stress decrease After grouting (kPa) 0.152 m square pile 1.83 1B 13.6 7.9 41.9% 130.3 0.203 m square pile 1.22 2.25B 12.5 10.4 16.8% 4 8.3 1.83 6.4B 17.4 17.2 1.1% 27.6 Table 3 2. Measured grout pressure and grout volume 0.203 m x 2.44 m pile 0.152 m x 2.44 m pile Grout pressure kPa (psi) Grout volume m 3 (Gal) Grout pressure kPa (psi) Grout volume m 3 (Gal) Top zone grouting 5 50 (80) 0.1893 (50) 550 (80) 0.0946 (25) Bottom zone grouting 825 (120) 0.1893 (50) 825 (120) 0.0946 (25) Tip grouting 1035 (150) 0.1136 (30) 1035 (150) 0.0378 (10) Table 3 3. Material properties used in PLAXIS Parameter Sand Pile E (kPa) 2.48 x 10 7 E 50 ref (kPa) 3.69 x 10 4 E oed ref (kPa) 3.69 x 10 4 E ur ref (kPa) 1.108 x 10 5 P ref (kPa) 100 unsat (kN/m 3 ) 16 25 sat (kN/m 3 ) 19 Friction angle, 31 Dilation angle, 0 0.25 0.15 Power, m 0.5
86 Table 3 4. Measured and predicted grout pressures for 2.44 m long piles Side grouting Tip grouting top bottom Measured Pressure(kPa) 550 825 1035 385 720 1450 445 760 1620 FEM (kPa) 345 620 1200 PMT (kPa) 415 520 -Table 3 5 Side resistance prediction Pile width (m) Pile length H(m) Initial vertical eff. stress at H/2 vo (kPa) K g at H/2 (Fig. 3 11 ) Grouted v ert. eff. stress vg =K g vo at H/2 (kPa) c f s (kPa) (Eq. 3 2 ) Bulb D iameter D f (m) Surface A rea A s (m 2 ) Side R esistance Qs (kN) 0.153 2.44 23.7 2.15 51 31 o 46.4 0.38 2.9 135 0.203 2.44 23.7 2.15 51 31 o 46.4 0.51 3.91 180.8 0.406 6.1 52.7 1.5 79 31 o 68 0.914 17.52 1191.4
87 Table 3 6 Estimation of require d jet pipe diameter Pile width D (cm) Pile length l (m) Soil Type Grain size d 50 (mm) Permeability k (m/day) Discharge Q (m 3 /hr) Velocity V (m/s) Jet pipe area A = Q/V (mm 2 ) Jet P ipe Diameter (mm) 71.1 4.57 A3 0.17 a 11.23 b 82.3 c 5 d 4572 76.2 a from gr a in size analysis b for A3 soil, Smith and Bloomquist (2010) c using Eq uation (2 1) d Tsinker (1988) Table 3 7 Comparison of measured and predicted grout pressures Top membrane Bottom membrane Pile 1 Pile 2 Pile 1 Pile 2 Measured Maximum Pressu res (kPa) 690 830 620 690 966 1 103 1 240 1380 (1991) solution ( kPa ) a, b 497 497 960 960 kPa ) a, b 697 697 1200 1200 PMT ( kPa ) 780 c 560 c 1 366 d 1 020 d a c = 29 o = 0 o & = 0.3, Dr = 40 % b Corresponds to the middle of top membrane = 1.98 m and middle of bottom membrane = 4.27 m c At a depth of 2.6 m d At a depth of 4.9 m
88 Table 3 8 Comparison of the measured and predicted tip grout pressures Pile 1 Pile 2 Me asured tip grout pressure (kPa) 2000 2206 2193 2193 2 900 2 900 Table 3 9 Limiting velocity suggested by AASHTO Designation R8 81 Type of situation Limiting velocity, mm /s (in/s) Historical sites or other critical locations 2.54 (0.1) Residential buildings, plastered walls 5.08 7.6 (0.2 0.3) Residential building in good repair with gypsum board walls 10.16 12.7 (0.4 0.5) Engineered structures, without plaster 25 .4 28.1 (1.0 1.5) Source: AASHTO Designation R8 81 Table 3 10 Side resistance prediction for field jetted and grouted pile Grout zone Depth to center, H (m) vo at H ( kPa ) K g at H (Fig 3 11 ) vg =K g vo ( kPa ) c f s ( kPa ) ( Eq. 3 2 ) Bulb D iamete r ( m ) Surface A rea ( m 2 ) Side Res istance ( kN ) Top 1.98 33 1.8 59.4 29 o 49 1.1634 7.80 382.2 Bottom 4.27 62 1.35 83.7 29 o 69 1.1683 7.83 540.3
89 Table 3 11 C omparison of measured and predicted side resistance Method Top membrane Bottom membrane Total M easured (kN)* 431 823 1254 Proposed methodology (kN) 382.2 540.3 922.5 Tip grout data (kN) --1589 *not fully mobilized Table 3 1 2 Comparison of unit skin frictions for jet ted and grouted pile vs. drilled shaft Segment Unit skin (kPa) Jet ted and grouted pile Drilled shaft Top membrane (0.91 3.05 m) 50.0 32.1 Bottom membrane (3.2 5.3 m) 98.2 37.5 Table 3 13. Forces and moments on the foundation for the E7 T6 Mast Arm assembly (design wind speed = 130 mph) Forces and moments E7 T6 type a New M ast arm b Torsion, M y (kN m) 351 352 Moment about axis of arm, M x (kN m) 202 201 Moment about axis normal to arm, M z (kN m) 158 160 Lateral load, V x (kN) 1.3 0 Lateral load, V z (kN) 33 33 Axial load, V y (kN) 25 47.6 a at design wind speed = 130 mph b at point lateral load = 33 kN Table 3 14. Dimensions of Mast arm assembly Length m (ft) Diameter cm (in) Thickness cm (in) Taper angle (deg ree ) Arm 12.2 ( 40 ) 50.8 ( 20 ) 1.6 ( 0.625 ) 0 Pole 6.1 ( 2 0) 61.0 ( 24 ) 1.6 ( 0.625 ) 0
90 Table 3 1 5 Force s and moments on the pile under maximum lateral load (54 kN) Forces and moments Magnitude Torsion, M y (kN m) 577.6 Moment about axis of arm, M x (kN m) 330 Moment about axis normal to arm, M z (kN m) 160 Lateral load, V x (kN) 0 Lateral load, V z (kN) 54 Axial load, V y (kN) 47.6 Table 3 16. Comparison of mobilized and predicted (ultimate) torsional resistance Method Magnitude Measured (kN m)* 577.6 Proposed methodology (kN m) 500 Tip grout data (kN m) 896 *not fully mobilized
91 Figure 3 1. FDO (Photo courtesy of author, Sudheesh Thiyyakkandi) Figure 3 2. Earth pressure cells in the test chamber (Photo courtesy of author, Sudheesh Thiyyakkandi)
92 Figure 3 3. Grain size dist ribution of test soil Figure 3 4. E xpansion pressure vs. volume curves from Pressuremeter tests 0 10 20 30 40 50 60 70 80 90 100 0.01 0.1 1 10 Percentage Passing (%) Grain Size (mm) 0 100 200 300 400 500 600 0 10 20 30 40 50 60 70 80 90 100 Pressure (kPa) Volume (cm3) Depth = 0.91 m Depth = 1.83 m
93 Figure 3 5. Residual horizontal stress variation with time near grouted pile Figure 3 6. Finite element discretization 0 25 50 75 100 125 150 175 200 0.01 0.1 1 10 100 residual horizontal stress after grouting (kPa) no. of days after grouting Depth = 1.83 m, 0.152 m away from 0.152 m pile Depth = 2.36 m, 0.152 m away from 0.152 m pile Depth = 1.83 m, 1.3 m away from 0.203 m pile Pile High resolution mesh close to pile 2.44 m (96 in) Elastic zone (Loose grout) Sand Axis of s ymmetry Chamber wall 1.83 m (72 in) 5.08 m (200 in)
94 Figure 3 7. FE mesh after the simulation of J et ted and grouted pile installation Figure 3 8 Stress distribution with radial distance during tip grouting (spherical cavity expansion) 0 200 400 600 800 1000 1200 1400 1 2 3 4 5 6 7 radial stress (kPa) Non dimensional radial distance (r/a) Analytical (Yu and Houlsby 1991) Measured radial stress Measured grout pressure FEM
95 Figure 3 9. Radial displacement determined from the numerical simulation of side and tip grouting of 0.203 m square p ile 0 20 40 60 80 100 120 140 160 180 0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2 Radial displacement (mm) Radial distance from the pile axis, m Tip grouting Side grouting of lower zone
96 Figure 3 10. ing min = vg N Shear Stress, f s Failure shear stress F ailure Stress State, Failure Pole Failure Plane Failure Normal stress max c
97 Figure 3 11. Estimate of grout vertical stress coefficient, K g 0 1 2 3 4 5 6 7 8 9 10 0.5 1.0 1.5 2.0 2.5 3.0 3.5 Depth (m) K g (dimensionless) FE analysis FE analysis FE analysis From pressure cell data Back calculated from pile's unit skin friction c = 31 o 33 o 35 o ( c = 31 o ) ( c = 33 o ) ( c = 35 o )
98 Figure 3 12. Comparison of load displacement curves for 0.152 m square pile Figure 3 13. Comparison of load displacement curves for 0. 203 m square pile 0 50 100 150 200 250 300 350 0 5 10 15 20 25 30 Load (kN) Displacement (mm) Experimental Proposed methodology FEM 0 100 200 300 400 500 600 0 5 10 15 20 25 30 35 Load (kN) Displacement (mm) Experimental Proposed methodology FEM
99 Figure 3 14. Comparison of load displacement curves for 0.406 m square pile Figure 3 15. L ayo ut of test piles and drilled shaft alo ng with reaction drilled shafts 0 500 1000 1500 2000 2500 3000 3500 4000 4500 0 5 10 15 20 25 30 35 Load (kN) Displacement (mm) Experimental (Lai et al. 2010) Proposed methodology FEM 11 m (36 ft) 11 m (36 ft) 11 m (36 ft) Reaction drilled shafts Test drilled shafts Jetted and grouted piles
100 Figure 3 16. SPT blow count (N) profiles and the Unified Soil Classification (USC) at the location of test piles and drilled shaft
101 Figure 3 17. Typical grain size distribut ions for the silty sand at the site Figure 3 18. Expansion pressure volume curves from Pressuremeter tests 0 10 20 30 40 50 60 70 80 90 100 0.01 0.1 1 10 Percentage passing (%) Grain size (mm) 0 200 400 600 800 1000 1200 1400 1600 0 20 40 60 80 100 Pressure (kPa) volume (cm 3 ) Pile 1: 2.6 m depth Pile 2: 2.6 m depth Pile 1: 4.9 m depth Pile 2: 4.9 m depth
102 Figure 3 19 Schematic diagram of jetted and grouted pile Strain gages
103 Figure 3 2 0 Grout deliv ery and jetting systems Figure 3 2 1 Reinforcing cage with grout delivery and jetting systems (Photo courtesy of author, Sudheesh Thiyyakkandi) Top Middle Bottom Strain gage Grout delivery system Jetting system
104 Figure 3 2 2 Concrete placement for one of the precast piles (Photo courtesy of author, Sudheesh Thiyyakkandi) Figure 3 2 3 Preparation of pile for jetting (Photo s courtesy of author, Sudheesh Thiyyakkandi) a) Flushing each grout delivery systems b) Attaching side grout membranes d) Testing of nozzles c) Attaching nozzles
105 Figure 3 2 4 Pile jetting (Photo courtesy of author, Sudheesh T hiyyakkandi) Figure 3 2 5 Schematic of precast concrete cap pile connection High pressure jet pump Water tanker Hydraulic pump for water recirculation Temporary casing
106 Figure 3 2 6 L ongitudinal and cross section view of concrete cap with reinforcement details
107 Figure 3 2 7 Placement and grouting of precast cap (Photos courtesy of author, James F Stephenson III) Figure 3 28 Casting of cast in place concrete cap (Photo courtesy of author, James F Stephenson III)
108 Figure 3 29 Side grouting of piles (Photo courtesy of author Sudheesh Thiyyakkandi) Figure 3 30. Grout pressure volume response during side grouting of pile 1 0 200 400 600 800 1000 1200 1400 0 0.2 0.4 0.6 0.8 1 1.2 1.4 Grout pressure (kPa) Grout volume (m 3 ) Pile 1 top bag Pile 1 bottom bag Grout mixing and pump ing Pile displacement and grout pressure monitoring
109 Figure 3 31. Grout pressure volume response during side grouting of pile 2 Figure 3 32. Grout pressure volume response during tip grouting 0 200 400 600 800 1000 1200 1400 1600 0 0.2 0.4 0.6 0.8 1 1.2 1.4 Grout pressure (kPa) Grout volume (m 3 ) Pile 2 top bag Pile 2 bottom bag 0 500 1000 1500 2000 2500 3000 0 0.1 0.2 0.3 0.4 0.5 0.6 Grout pressure (kPa) Grout volume (m 3 ) Pile 1 Pile 2
110 Figure 3 33. Grout pressure vs. pile head displacement during tip grouting Figure 3 34. Instrumentation for noise and vibration measurement (Photos courtesy of author, Sudheesh Thiyyakkandi) 0 500 1000 1500 2000 2500 3000 0 2 4 6 8 10 12 14 16 Grout pressure (kPa) Pile top displacement (mm) Pile 1 Pile 2 Triaxial geophone Recorder Microphone
111 Figure 3 3 5 Location of construction equipments and vibration and noise monitors
112 Figure 3 36. Noise measurement during pile jetting process Figure 3 37. Noise measurement during side grouting 40 50 60 70 80 90 100 110 120 0 2 4 6 8 10 12 14 16 18 20 22 24 26 28 30 Noise (dBA) Distance from pile (m) Pile1_during pile jetting Pile2_during pile jetting Impact pile driving (FHWA_RCNM)
113 Figure 3 38. Noise measurement during tip grouting Figure 3 3 9 Peak particle velocity measurement during pile jetting Figure 3 40 Peak particle velocity measurement during grouting 0 1 2 3 4 5 6 0 5 10 15 20 25 30 Peak particle velocity (mm/s) Distance from pile (m) Pile 1_during pile jetting Pile 2_during pile jetting
114 Figure 3 41. Axial top do wn test on jetted and grouted pile (Photo courtesy of author, Sudheesh Thiyyakkandi) Wooden reference beam Load cell Dial gage Mirrored scale Hydraulic jack Invar staff
115 Figure 3 42. Measured strain at different depths Figure 3 4 3 Load displacement response of the jetted and grouted pile 0 200 400 600 800 1000 1200 1400 1600 1800 0 1 2 3 4 5 Load (kN) Displacement (mm) Jet grouted pile Top load Jet grouted pile Tip load Jet grouted pile Skin resistance
116 Figure 3 44. Load distribution along the pile
117 Figure 3 4 5 Lowering of reinforcing cage and concrete placement for test drilled shaft (Photos courtesy of author, Sudheesh Thiyyakkandi) Figure 3 4 6 Load dist ribution along the drilled shaft Drilled shaft
118 Figure 3 4 7 Load displacement response of drilled shaft Figure 3 4 8 Comparison of axial response of jetted and grouted pile vs. drilled shaft 0 200 400 600 800 1000 1200 1400 1600 0 2 4 6 8 10 12 14 16 Load (kN) Displacement (mm) Top load Tip load Skin resistance 0 200 400 600 800 1000 1200 1400 1600 0 1 2 3 4 5 Load (kN) Displacement (mm) Jetted and grouted pile Drilled shaft
119 Figure 3 49 Coordinate system used for representing forces and moments Figure 3 50 Mast arm assembly fabricated for combined torsion and lateral load test (Photo courtesy of author, Sudheesh Thiyyakkandi) X Z Y Axis parallel to arm Axis of pile Pile
120 Figure 3 5 1 Setting and bolting Mast arm assembly ( Photos courtesy of author, Sudheesh Thiyyakkandi) Bolt ing the pole flange to pile cap Setting Mast arm assembly using a crane
121 Figure 3 5 2 Total Station for rotation and translation measurement (Photo courtesy of author, Sudheesh Thiyyakkandi) Total station P ole Top view Axis of arm Targets on pole
122 Figure 3 5 3 String pot arrangement (Photo courtesy of author, Sudheesh Thiyyakkandi) 4 String pots in each level (total: 2 x 4 = 8) Mast arm Side view
123 Figure 3 5 4 Digital dial gage placement (Photo courtesy of author, Sudheesh Thiyyakkandi) Dial gages Wooden frame Pole
124 Figure 3 5 5 Application of lateral load on the arm by pulling wi th a crane (Photo courtesy of author, Sudheesh Thiyyakkandi) Figure 3 5 6 Tension load cell for the load measurement (Photo courtesy of author, Sudheesh Thiyyakkandi)
125 Figure 3 5 7 Torque vs. rotation response during combined torsion and lateral load test Figure 3 5 8 Lateral displacement components during the load test 0 100 200 300 400 500 600 700 0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2 Torque (kN m) Rotation (degree) Using Total station data Using string pot data Using Dial gage data x y Mast arm
126 Figure 3 5 9 Lateral load vs. r esultant lateral displacement during combined torsion and lateral load test 0 10 20 30 40 50 60 0 2 4 6 8 10 12 14 16 18 20 Lateral load (kN) Resultant lateral displacement (mm) Using Total station data Using Dial gage data
127 Figure 3 60 Torsional cracks and gaps after loading and unloading (Photos courtesy of author, Sudheesh Thiyyakkandi) Aft er loading Torsional cracks Gap (0.75 in) After un loading Gap (0.125 in) Torsional crack
128 CHAPTER 4 GROUP RESPONSE OF JETTED AND GROUTED PILES This C hapter presents the experimental study carried out to investigate the group behavior of jet t ed and grouted piles at typical 3D center to center spacing The tests 1) at the Coastal Engineering Lab, University of Florida Past studies have shown that the boundary conditions and size of a test chamber may influence t he soil stress and displacement field around pile, tip resistance of a penetrating object (e.g., cone penetrometer) and limit pressure of an expanding object ( Meyerhof 1959, Kishida 1963, Robinsky and Morrison 1964; Fahey 19 80 ; Schnaid and Houlsby 1991 ; Salgado et al. 1998 ; Kraft 1991 ). The zone of influence around a pile during installation or axial loading is found to be varied from 3 to 8 pile diameters depending on the soil density and installation method (Meyerhof 1959, Kishida 1963, Robinsky and Morrison 1964). Randolph and Wroth (1978) stated that under axial loading, the influence zone around pile will extent radially about 0.9 modulus distribution wit h depth. According to Kulhawy et al. (1979), the boundary influence is negligible when the chamber to pile diameter ratio is greater than 14. Kraft (1991) reported that the zone of influence reaches the wall of the chamber when pile length to diameter rati o is equal to or greater than 0.45 times the chamber to pile diameter ratio. For a given chamber size, the boundary effect seems to be more significant for dense sand than loose to medium sand ( Schnaid and Houlsby 1991 ). In view of that, it was decided to conduct the group tests in loose to medium dense state. The size of the piles within the group was decided by considering boundary constraints of the testing chamber and cylindrical cavity expansion limit pressure. Based
129 on the cavity expansion formulatio n of Yu and Houlsby (1991), a cylindrical expansion ratio (pile radius to grout zone radius) of 2 is at least required to develop an expansion pressure of about 80 85% of the limit pressure in the present soil state. To maximize the capacity of jetted and grouted pile, the pile need to be grouted to the highest possible grout pressures, which is equal to the cavity expansion limit pressures (McVay et al. 2009, Thiyyakkandi et al. 2012). Two groups were considered in the study; the first group consisted of f our 20.3 cm square 2.44 m long (8 in square x 8 ft. long) precast piles with 40.6 cm (16 in) diameter side grout bulbs at 61 cm (3 x pile width) center to center spacing; the second group was four 11.4 cm (4.5 in) diameter 2.44 m piles with 2 5.4 cm (10 in) diameter side grout bulbs For the first group, the chamber boundary was only 3.5 x bulb diameters (or 7 x precast pile widths) away from the piles, and hence, the group capacity might be influenced by the boundary. However, for the second group, the chamber boundary was about 6.25 x bulb diameters (or 15 x precast pile diameters) away from the piles and hence will have lesser boundary influence. A description of construction and testing of both groups is presented followed by an analysis of the result s 4.1 Group Testing of Jetted and Grouted Pile s 4.1.1 Soil Preparation in the Test C hamber Typical Florida silty sand (A 2 4) used in the previous test chamber study of jetted and grouted piles ( S ection 3.1) reported by McVay et al. (2009) was used as the test so il in the group study. For all the tests, moisture content of the soil was in the range of 5 to 9% The test chamber was filled in layers (lifts) and the average thickness of each lift was 0.46 m (1.5 ft). After placing the adequate quantity of soil throug h free
130 fall from the top of the chamber, each lift was compacted uniformly using a vibratory plate compactor (Figure 4 1) Soil dry density for all the lifts measured using core cutter method/nuclear density gage were in the range of 15.56 16.25 kN/m 3 (99 103.5 lb/ft 3 ). Using the dry density values, the relative density of compacted soil layers for all the tests were estimated to be in the range of 34 54%. While filling the test chamber, a number of earth pressure cells were placed vertically and horizon tally at various locations to measure the vertical and radial soil stress changes during the installation and testing of piles/ shafts (Figure 4 2) 4.1.2 Design and Construction of Precast P iles The precast piles for the first group were traditional reinforced concrete component with square cross section (20.3 cm x 20.3 cm). Whereas the structural section of the second group piles was 11.4 cm (4.5 in) steel pipe (6.35 mm thick) filled with concrete. Size of grout distribution systems and jetting systems for the piles were selected in conformity with Thiyyakkandi et al. (2012). The jetting system for square piles (group 1) was made using 3.8 cm (1.5 in) steel pipes and that for 11.4 cm diameter piles (group 2) were made using 2.5 cm steel pipes. The jetting syste m was also used for tip grout ing during the grouting phase. Two separate grout delivery systems were designed for the top a nd bottom halves of each pile. In the case of 20.3 cm square piles, side grout delivery systems made of 2.54 cm PVC pipes were embedd ed within the pile section, running a long the corners of the piles. While the side grout delivery systems for 11.4 cm diameter piles were located outside the pile (welded to the pipe) due to the limited space available inside the piles. Each of the grout p ipes (entry and exit) has a series of holes drilled into the PVC/steel pipes in pairs at 10 cm intervals. A 2.5 cm diameter gum rubber membrane (6.4 mm thic k) covered each pair of holes.
131 Fig ures 4 3 and 4 4 shows the placement of the grout delivery systems and jetting/tip grouting pipes for the precast piles of first and second group respectively After sufficient curing following casting of piles, water under pressure was flushed through each grout system to ensure that the grout systems (i.e., top and bot tom) were working properly. Subsequently, semi rigid membranes were attached to the upper and lower portions of each pile as shown in Figure 4 5 Rubber nozzles were then attached to the bottom of each jet pipe to reduce the cross sectional area of the pip e at tip (Figure 4 5 ) 4.1.3 J etting of Precast Piles for Each Group For each group, jetting of the piles into the test chamber was done one by one at a time interval of 24 hours. Figure 4 6 shows the jetting of one of the group 1 precast piles into the test cha mber. A sufficient time interval (24 hours) was used between the jetting to allow water to percolate to a greater depth, and thus eliminate the quick condition created in the chamber due to the previous jetting. The water for jetting was supplied at flow r ate of 0.25 m 3 /minute and a pressure of 415 kPa. The pile s generally penetrated the soil under its own weight, but were guided to its final group position by hand. Approximately 4 to 5 minutes of jetting was required to install the pile s to the required d epth. Any water accumulated at the top of the chamber was pumped out and the water level in the test chamber was continuously monitored and pumped out of the PVC pipes located against the chamber wall to maintain a constant water table 4.1.4 Side G routing of t he Piles After jetting of all the piles for each group and ensuring that the water level in the test chamber was well below the pile tip, side grouting of the piles was initiated (Figure 4 7 ). Top zon e of all the piles in a group was grouted first. Groutin g of the bottom zones were performed after hydration of the top grout zones. Volume of grout pumped and
132 maximum sustained grout pressure observed for both top and bottom bag grouting of all the piles are given in Table 4 1. 4.1.5 Static Top D own Test Prior to Ti p Grouting Static top down test was performed on the side grouted pile groups (prior to tip grouting) to separate skin resistance of pile group from the total axial capacity. Figure 4 8 displays the test setup for the group load test. As shown from the Fig ure 4 8 the test setup consisted of load cells of 890 kN (200 kips) capacity on each pile to measure load distribution, a test frame, a hydraulic jack, another load cell of 2669 kN (600 kips) capacity to measure total load, a reaction beam with a support system, and the displacement monitoring instrumentation. In case of group 1, the vertical displacement of individual piles was monitored using digital levels and invar staffs, Figure 4 9. While for group 2, pile displacement was measured using digital dial gages. Soil deformations at the center of group and at different radial distances towards the chamber were also monitored using digital dial gages to investigate group effect. The tests were continued until the full mobilization of skin resistance, which was ensured from the log log plot of load versus displacement for each of the shafts. The applied load corresponding to the change in slope on the log log plot was considered as the ultimate skin resistance of each side grouted pile because the lo ad transferred through pile tip prior to the slope change wa s insignificant as identified from the earth pressure cells installed directly below piles tips (negligible soil stress change). 4.1.6 Tip G routing of the Piles After the top down compression test of t he groups tip grouting of each pile in a group was carried out individually Grouting of each pile was stopped when the top of each pile moved upward approximately by 6.35mm (i.e., full mobilization of side
133 resistance). Tip grouting data of piles in each group is also given in Table 4 1. Note that the observed maximum sustained grout pressures and grout volumes were in the same range for all the piles in each group. 4.1.7 Static Top D own Test After Tip Grouting Following tip grouting, top down load testing was conducted on each group to esti mate the total axial capacity. The load test setup was the same as that for the group tests prior to tip grouting. Besides the load displacement response of piles, soil deformations in the vicinity of group were also monito red during group loading. 4.1.8 Excavation of Jetted and Grouted Pile Groups Shown in Figure 4 10 are the jetted and grouted pile group s after excavation. Each pile in the group appears to possess good quality grout zones In case of group 1, the diameter of the up per side grout bulb varied from 27.5 to 47.5 cm along the pile, and the diameter of the lower side grout bulb varied from 30 to 45 cm along the pile. The diameter of all side grout bulbs for group 2 was in the range of 24 26 cm as expected. Tip grout bu lbs were nearly spherical (Figure 4 10 and 4 11) with diameter in the range of 37 48 cm for group 1 and 28 30 cm for group 2. 4.2 Analysis of Experimental Jetted and Grouted Pile Group Behavior This section analyz es the result s obtained from the experimental study and discusses the group behavior of the piles. Table 4 2 presents the highest horizontal stress change measured near chamber boundary for group 1 during grouting and group load test The maximum soil deformations observed near chamber boundary durin g group load tests were 3.3 mm and 0.64 mm for group 1 and 2 respectively. These soil stress and deformation measurements indicate that there was some boundary ef fect in the case of Pile group 1 For group 2 the measured soil deformation near the chamber
134 wall was negligible ( 0.64 mm) compared to group 1, which is due to the fact that the chamber boundary for group 2 was 6.25 x bulb diameters away from the piles ( approximately double of group1 ) Therefore, the boundary effect in group 2 was assumed to be in significant Figure 4 12 presents the total load versus average displacement of pile group 1 and individual response of each pile in the group during the load test performed prior to tip grouting. Even though the distribution o f applied load was non unifo rm as shown in Figure 4 12 the vertical displacements of piles were relatively uniform. Similar response was observed for group 2 also. Moreover, for both groups, the soil deformation at the center of group was nearly same as the average displacement at t he pile heads for each load increment. These observations suggested that the load has been transferred through shear within the group and the group a cted as a single block during. By plotting load displacement response of groups in log method), it was found that skin resistance of group 1 was at least 302 kN ( 68 kips ) and that of group 2 was approximately 196 kN ( 44 kips ) Vertical stress measurement beneath pile group lot, which suggests that the entire applied load was carried as skin resistance. Shown in Figures 4 1 3 and 4 1 4 are the loads displacement s curves from the top down testing of group 1 and 2 respectively after tip grouting. It is evident from the shape o f the curves that significant end bearing was mobilized for the group response Ground surface crack observed around group during axial loading as shown in F igure 4 1 5 indicates that the shearing occurred around group but not around individual piles For b oth groups the displacements of piles and the soil deformation within the group
135 were relatively uniform during loading ( Figures 4 13 4 14 4 16 and 4 17 ). These observations suggest the block behavior of the g roups during top down loading. Outside the gr oups, soil deformation was found to diminish quadratically with increase in distance from the group boundary as displayed in Figures 4 1 6 and 4 1 7 Horizontal stress measurement around piles during the axial load test ( Figure 4 1 8 ) indicated that the stres s decreased with an increasing axial load until the full mobilization of skin resistance (5 load steps), and then, increased wit h an increase in applied load. During the unloading phase ( Figure 4 1 8 ), the la teral stress always decreased. It is believed tha t the decrease in lateral stress during skin resistance mobiliza tion was due to the rotation of the origin of planes or pole such that the failure plane aligns in the vertical direction (Thiyyakkandi et al. 2012). Shown in Figure 4 19 is the variation of v ertical stress below (0.61 m) the center of the group footprint and one of the piles during the top down load test ing of group 1 before tip grouting Similarly, Figure 4 20 shows the same during the load test after tip grouting. Of interest is the increase in stress at the center of the pile group versus the edge (West pile) in both cases supporting the superposition (overlap) of stresses from the adjacent piles Similar stress variation was observed in group 2 load test s also. The experimental study of je tted and grouted pile groups suggested a significant increase in both densification of soil around the piles with an increase in the horizontal h ) and shear modulus of soil in close proximity of the pile s C onsequently, the pile has a much higher ultimate side resistance compared to tradition al driven piles/drilled shafts. In the case of a jetted and grouted piles installed in a group, the side grouting of adjacent piles increased the confining stress and shear modulus of the soil
136 confined within the group. So the soil within the footprint of the group underwent much smaller shear strains and the pile soil mass behaved as a rigid body. However, outside the footprint of the jet ted and grouted pile group, the horizontal (radial) stresses Therefore, the grouted pile group failed as a block (i.e ., uniform movement) within the group, but quadratically outside the group under axial loading. The vertical side capacity of the group may be obtained by multiplying the shear stress on the surface of any single jet grouted pile with the surface area of t he block (i.e., block perimeter of the group the length of the pile). In the case of the group tip resistance, grouting of the individual piles developed spherical cavity expansion stresses around each pile, as well as densification of soil (i.e., increa sed shear modulus) between the bulbs (primarily within the group foot print) through displacement of the soil in the radial direction. Consequently, when the group was loaded, the stresses from each pile transferred to the center of the group (i.e., superp osition) resulting in higher stresses at the pile group center (Figure s 4 19 and 4 20 ) and block tip resistance (group foot print area unit tip resistance) under axial top down loading developed. Figure 4 2 1 depicts the comparison of group response befo re and after tip grouting of group 1 A higher initial slope of curve in load test 2 (grouped pile tips) versus test 1 (grouted pile sides only) is due to initial mobilization of the end bearing during the mobilization of skin resistance.
137 4.3 Predicted Axial R esponse of Jetted and Grouted Pile Groups 4.3.1 Skin Resistance of the Groups Since the piles act as a block during axial loading, the total side resistance of the group can be estimated by multiplying the unit skin friction with the surface area of the bl ock. U nit skin resistance of a single jet ted and grouted pile can be estimated using the E quation 3 2. Table 4 3 presents the prediction of the average ultimate unit skin friction ( f s ) and side resistance ( Qs ) of the single jet ted and grouted pile s and the group s u sing Equations 3 2 and 3 3, and Figure 3 11 The unit skin friction at a depth equal to the embedment depth of the piles was first determined using K g plot ( Figure 3 11 ) and unit skin friction equation (Equation 3 2). Then, the average skin friction alo ng the pile was determined by the assuming a linear variation of skin friction with depth. Side resistance of the individual piles and the groups were obtained by multiplying the average unit skin friction with the pile surface area and the block surface a rea respectively. It can be seen from the Table 4 3 that that the group shear resistance is less than the sum of individual shear resistances, which is attributed to the reduced block surface area of the group versus the sum of the individual pile surface areas Tip grout pressure can also be used to estimate skin r esistance of individual piles. The ultimate axial skin friction of a pile should be equal to the tip grout pressur e times an effective tip area. In case of square pile, effective area is assumed as the area of circle with diameter equal to the diagona l distance of pre cast pile. Whereas for the circular pile, the effective tip area used in the calculation was the area of the bottom plate (Figure 4 5 ) used to hold bottom end of side grout membrane. Table 4 4 provides a comparison of the skin resistance of single pile predicted using the tip grout pressure
138 with that predicted by the proposed method (Equation 3 2) and is evident that the predictions matched reasonably well 4.3.2 Load Displacement Respons e of the Groups The load displacement prediction approach suggested for single jetted and grouted piles described in Section 3.3.3 was used here to predict the load displacement response of the groups Table 4 5 lists the various parameters used for the pr ediction. Note that r 0 and R 0 values used in the prediction w as the equivalent radius of the block (i.e., equivalent radius of group footprint) Shear modulus of the soil was obtained using the equation suggested by Randolph et al. (1994) for pile design: (4 1) Where p a is atmospheric pressure (100 kPa), is the mean effective stress, and S is a constant that decreases from 400 to 75 as the silt content increases from 5 % to 15 % ( Randolph et al. 1994) The surface area of the block was used for estimating total side resistance of the group and the effective block footprint area was used for estimating ultimate tip resistance ( Q f ). Total axial load displacement response of groups was then obta ined from the combination of the T Z response and Q Z response. Shown in Figures 4 2 2 and 4 2 3 are the predicted and measured load displacement responses of the group 1 and group 2 respectively. The predicted group load displacement curves were in good agr eement with the measured load settlement response s
139 Table 4 1 J etted and grouted piles grouting data Group Pile ID Grout pressure, kPa Grout volume m 3 (gal) Top bag a Bottom bag a Tip Top bag Bottom bag Tip PG1 S 517 827 1517 0.090 (24) 0.094 ( 25) 0.079 (21) N 483 827 1724 0.088 (23.5) 0.094 (25) 0.064 (17) W 413 896 1655 0.086 (23) 0.094 (25) 0.079 (21) E 448 862 1793 0.090 (24) 0.094 (25) 0.071 (19) PG2 W 345 827 1689 0.043 (11.5) 0.043 (11.5) 0.019 (5) E 345 759 1724 0 .043 (11.5) 0.043 (11.5) 0.019 (5) S 358 1034 1862 0.043 (11.5) 0.043 (11.5) 0.023 (6) N 393 1034 2138 0.043 (11.5) 0.043 (11.5) 0.026 (7) a After deducting the pressure lose in grout distribution system (138 kPa) from the measured pump pressure Tab le 4 2 Highest horizontal soil stress increase near chamber boundary during grouting and group load test Stage Stress change Measurement point elevation Side grouting 54 kPa Middle of bottom bag Tip grouting 58 kPa Pile tip Group load test 24 kPa Pile tip
140 Table 4 3. Side resistance prediction for single piles and groups Group Initial vertical eff. stress at embedment depth ( H ), vo (kPa) K g at H (Fig. 3 11) Grouted vert. eff. stress at H vg =K g vo (kPa) c Unit skin at H f s H (kPa) (Eq. 3 2 ) Ave. skin a f s = f s H /2 (kPa) Bulb Diameter D f (m) Surface Area, A s (m 2 ) Side Resistance, Qs (kN) Single Group Single Group Group 1 42 1.67 70.14 31 o 63.85 31.925 0. 406 3.11 9.06 99.3 298.2 Group 2 42 1.67 70.14 31 o 63.85 31.925 0.254 1.946 5. 29 62.1 168.9 a Assuming linear variation of unit skin friction with dep t h Table 4 4 Comparison of individual skin resistance prediction using the tip grout pressure and the proposed method Group Ave. tip grout pressure q g (kPa) Effective tip area A ( c m 2 ) Individual pile skin resistance = q g .A /10 4 (kN) Predicted skin resistance using Eq. 3 2 (kN) Group 1 1672 648.4 108.4 99.3 Group 2 1853 366.1 67.8 62.1
141 Table 4 5 List of parameters used for the group load displacement prediction Paramet ers Group 1 Group 2 r 0 (cm) 57.4 33 R 0 (cm) 57.4 33 L (cm) 2.44 2.44 0.25 0.25 G L/2 (kPa) a 12531 12531 G tip (kPa) a 17734 17734 max b 31.925 31.925 P L (kP a ) c 1450 1450 R f 0.99 0.99 R t 0.99 0.99 Block tip area, A block (m 2 ) 0.9968 0.3424 a Using Equation 4 1 b Using Equation 3 2 c Using the solution of Yu and Houlsby (1991)
142 Figure 4 1. Soil compaction using vibratory plate compactor (Photo courtesy of author, Sudheesh Thiyyakkandi) Figure 4 2 Earth pressure cells placement (Photo courtesy of author, Sudheesh Thiyyakkandi)
143 Figure 4 3 Reinforcing cages with grout delivery and jetting systems for group 1 piles (Photos courtesy of author, Sudheesh Thiyyakkandi)
144 Figure 4 4 Reinforcing cages with grout delivery and jetting syste ms for group 2 piles (Photos courtesy of author, Sudheesh Thiyyakkandi)
145 Figure 4 5 Attachment of semi rigid membranes and typical rubber nozzle used (Photos courtesy of author, Sudheesh Thiyyakkandi)
146 Figure 4 6 Pile during jetting (Photo courtesy of author, Sudheesh Thiyyakkandi) Figure 4 7 Side grouting of pile in group 2 (Photo courtesy of author, Sudheesh Thiyyakkandi)
147 Figure 4 8 Load test setup for jetted and grouted pile group (Photo courtesy of author, Sudheesh Thiyyakkandi) Fi gure 4 9 Digital levels and invar staffs used for pile displacement monitoring (Photos courtesy of author, Sudheesh Thiyyakkandi)
148 Figure 4 10 Jetted and grouted pile groups after excavation (Photos courtesy of author, Sudheesh Thiyyakkandi) Figu re 4 11 Views of a tip grout bulb (group 1) (Photos courtesy of author, Sudheesh Thiyyakkandi)
149 Figure 4 1 2 Load displacement response of group 1 prior to tip grouting Figure 4 13 Load displacement response of group 1 after tip grouting 0 50 100 150 200 250 300 350 0 1 2 3 4 5 Load (kN) Displacement (mm) North Pile South Pile West Pile East Pile Total load vs ave. pile displacement Total load vs soil deformation at group center
150 Figure 4 14 Load displacement response of group 2 after tip grouting Figure 4 1 5 Ground surface crack around group during axial loading (Photo courtesy of author, Sudheesh Thiyyakkandi)
151 Figure 4 16. Soil deformation profile (group 1 load test) Figure 4 17. Soil deformation profile (group 2 load test) Group boundary Group boundary
152 Figure 4 18. Typical variation of horizontal stress around pile during axial load te st Figure 4 19. V ertical stress variation below the center of the group footprint v s. beneath pile during the top down load test before tip grouting 0 5 10 15 20 25 30 35 40 45 0 50 100 150 200 Horizontal stress (kPa) Time (min) North (1.22 m depth 0.46 m away from pile) East (1.22 m depth 0.46 m away from pile) South (1.22 m depth 0.46 m away from pile) West (1.22 m depth 0.46 m away from pile) 0 10 20 30 40 50 60 0 50 100 150 200 250 Vertical stress (kPa) Time (min) Below center of group footprint Below west pile
153 Figure 4 20 Typical vertical stress variation below the center of the group footprint v s. benea th pile during the top down load test after tip grouting Figure 4 21. Comparison of group response before and after tip grouting of group 1 0 100 200 300 400 500 600 700 800 0 50 100 150 200 250 Vertical stress (kPa) Time (min) Below center of group footprint Below west pile 0 200 400 600 800 1000 1200 0 10 20 30 40 50 60 Total load (kN) Average pile displacement (mm) Group test before tip grouting Group test after tip grouting
154 Figure 4 22. Predicted and measured load displacement response of group 1 Figure 4 23. Predicted and measured load displacement response of group 2 0 200 400 600 800 1000 1200 1400 0 10 20 30 40 50 60 Load (kN) Displacement (mm) Predicted Measured 0 100 200 300 400 500 600 700 800 900 0 5 10 15 20 25 30 35 40 Load (kN) Displacement (mm) Predicted Measured
155 CHAPTER 5 GROUP RESPONSE OF POST GROUTED DRILLED SHAFTS This C hapter focuses on the experimental investigat ion of group behavior of grout tipped drilled shaft groups. To minimize the chamber wall effects, small diameter shafts (21.6 cm or 8.5 in ) were chosen for the group study of PGDS. Two different behavior at different embedment depths. The first group considered for the study was the group of four 0.216 m diameter x 2.44 m long ( 8 .5 in diameter 8 ft lon g ) drilled shaft (L/D ~ 11) at 3D c/c spacing. The objective of the first group test was to study the influence of preloading, and tip bulb area on the axial capacity of grout tipped drilled shaft as well as study the grout flow pattern, and estimate the ax ial group efficiency of tip grouted shafts at 3D spacing. The second group chosen for the study was the four 0.216 m diameter x 3.96 m long ( 8 .5 in diameter 13 ft long ) drilled shaft (L/D ~18) group at 3D c/c spacing The intent of the latter test was t o validate the results obtained from the first group test at greater embedment depths and investigate the effectiveness of staged grouting in increasing the capacity of post grouted drilled shafts In both group tests the chamber boundary was about 6.5 di ameters away from the shafts, which only satisfies some of the zone of influence criteria discussed in C hapter 4 Hence there may be small boundary effects during the installation and axial group tests, which could be quantified from the stress and deforma tion measurement near the test chamber wall. A description of various stages of the experimental investigation, the data measured, as well as analysis of the results, is presented.
156 5.1 Group Testing of Post Grouted Drilled Shafts 5.1.1 Soil Preparation for Group Te sts at Coastal Engineering lab, UF, was used for the g roup study of post grouted drilled shafts in cohesionless soils. Soil used in the study was the same soil (typical Florida silty sand) used in the individual and group study of jetted and grouted piles. The moisture content of the soil was in the range of 5 to 9% In order to eliminate the drilling and difficulties associated with installing the shaft, the temporary PVC casings at 3 x shaft diameter center to center spacing wer e positioned in the test chamber before soil placement This also ensured the verticality of the shafts and allowed the placement of earth pressure cells in close proximity to the shafts. Figure 5 1 presents the placement of PVC casings before fillin g the test chamber The method of soil placement was the same as that for the group testing of jetted and grouted piles. Core cutter method and nuclear density gage were used to determine the soil dry density for all the lifts and found to be in the range of 15. 7 16 kN/m3 (99 .9 102lb/ft3). Figure 5 2 shows the test chamber in fully filled state. 5.1.2 Fabricati on of the Re inforcing Cage and Tip Grout System for Drilled Shafts A rea of s teel r einforcement for the drilled shaft s was calculated according to building code requirements for structural concrete (ACI 318 0 8 ). The reinforcing cage consisted of five #5 rebars as longitudinal reinforcement and #3 rebars at 0.3 m spacing for shear reinforcement. Shown in Figure 5 3 is the reinforcing cage and grout distribution sy stem for the drilled shafts. The grout distribution system used for the study was a sleeve port apparatus (tube a Manchette) with a steel plate and a scuff ring (15 cm in diameter and 10 cm long) welded to the bottom of a reinforcement cage (Figure 5 3 ), which is typical
157 practice in the Southeastern United States (Dapp et al. 2006 ). The sleeve port apparatus consisted essentially of a pipe network at the shaft tip with pre drilled holes wrapped in gum rubber tubes. The principal advantages of the apparatu s were to facilitate grout distribution and the feasibility of cleaning, or removing the grout in the pipe network for later re grouting of the shaft, if necessary. The gum rubber tubes allowed the grout to exit the pipe network at minimum grout pressure, 103 138 kPa (15 20 psi), and prevented the loss of water when cleaning the pipes at smaller pressures. The minimum grout pressure was obtained by testing the grout system with pressurized water. The bottom of the scuff ring was covered with an u ltra s oft p olyurethane membrane, which protected the sleeve port apparatus during casting of the shaft. Construction of all the shafts in each group was performed in the same day To cast the shaft, the rebar cage with its grout distribution system was first lowered into the PVC casing, and the fluid concrete was placed in the casing. Subsequently, the PVC casing was pulled (Figure 5 4) using a forklift and the top of the shaft s was leveled. 5.1.3 Top Down Testing of Drilled Shaft Groups Prior to Tip Grouting Each group wa s subjected to static axial load test (ASTM 1143 07) before tip grouting to estimate initial skin resistance. Figure 5 5 display s the test setup which consisted of spacer disks on each shaft span grout delivery pipe, 890 kN ( 200 kip ) capacity load cells pl aced on the top of each shaft to measure load distribution, a test frame, hydraulic jack, another 2669 kN ( 600 kip ) capacity load cell to measure total load, a reaction beam and support system, and displacement monitor ing instrumentation Shaft and soil d isplacement were measured using digital dial gages. The test load was applied in increments and each load increment was maintained for a constant time interval of 5 minutes. This time interval was considered adequate to
158 assess the vertical displacement (mo stly immediate settlement) under each load increment because the soil used in the study was silty sand and the water table was kept well below the bottom of shaft. The tests were conducted up to full mobilization of the skin resistance, identified from the log displacement and corresponding to a displacement of approximately 6.35 mm (0.25 in). Note that the change in slope on the log load vs. log displacement plot ( full mobilization of skin friction) was visible because t he tip resistance was negligible due to the unfilled steel ring membrane system at the base of the shaft (prior to tip grouting). 5.1.4 Tip G routing of Drilled Shafts After top down group test the tip of each shaft was grouted individually (Figure 5 6 ) The gr out mix consisted of cement, 10% micro fine fly ash and water at a water/cement ratio of 0. 45 For group 1 shafts ( 2.44 m long ) grouting was performed in single stage and continued until an upward displacement of about 6.35mm (0.25 in) at shaft head Tip grouting started with the south shaft (SG1 S) and the displacement of the shaft was negligible up to a grout volume of 0.053 m 3 ( 14 gallons ) The shaft reached an upward displacement of 6.35 mm when 0.064 m 3 (17 gallon) was p umped to the tip of the shaft. The maximum grout pressure observed during the grouting was 448 kPa. However, during the grouting of the other three shafts, upward displacement of 6.35 mm occurred under very small volume of grout as shown in Table 5 1 Note that the maximum pump pressur e observed for all the shafts were same even though the volume of grout pumped were different (Table 5 1 ) In case of group 2 shafts (3.96 m long), grouting was performed in multiple stages with different colored grouts in each stage to identify grout flo w pattern during different stages. In the first stage of grouting, 0.023 m3 (6 gallon) of yellow colored grout
159 was pumped into each shaft. After hydration of the first stage grout, another 0.023 m3 (6 gallon) of red colored grout (second stage) was pumped into eac h shaft. When the third stage grouting (black grout) was attempted in the east (SG2 E) and south (SG2 S) shafts, grout pressures exceeding 4138 kPa (600 psi ) were recorded, however, no grout was pumped into the shafts This was attributed to the formation of a solid grout plug at the shaft tip due to the fi rst and second stage grouting. As the third stage grouting was controlled by upward displacement (about 6.35 mm) of shaft, grouting of the west (SG1 W) and north (SG1 N) shafts discontinued af ter pumping of about 0.011 m 3 (3 gal lon ) of grout. The grout pressures measured during the stage 2 and 3 grouting was about 15 to 35% higher than that during the stage 1 (Table 5 1) 5.1.5 Top Down Testing on Tip Grouted Drilled Shaft Groups After the hydration of grout, top down testing on the PGDS groups were conducted to assess the total load capacity of the groups and the group interaction The test setup was exactly the same as that for th e load test prior to grouting. In order to estimate the interaction be tween the shafts, individual loading of a couple of shafts in each group were also performed (Figure 5 7 ) Besides the load displacement response of test shaft, displacements of other shafts and stresses in the vicinity of the group were also measured duri ng the individual shaft load tests 5.1.6 Excavation of Tip Grouted Drilled Shaft Groups During the excavation of the grouted shaft groups it was observed that little, if any, grout bulb formed beneath the shaft bases. In addition, no grout permeation or grout soil mixing occurred below any of the tested shafts. Figure s 5 8 and 5 9 shows the photographs of the excavated group 1 and group 2 shafts respectively and it is evident that the grout flowed up along the shaft soil interface during grouting every time.
160 T he upward migration of the grout was attributed to the minor principal stress acting in the horizontal direction along the shaft soil interface, making the interface a major conduit for grout flow. The fluid grout also relieved any side friction (viscous f luid) when it flowed upward, as has been observed in a case study in Walton County, Florida (Muchard 2004) However, the upward after hydration, which in turn would increase the skin resistance after hydrati on. For instance, the bottom radius of group 1east shaft (SG1 E) increased by 33 mm (1.30 in, 15 %) due to upward flowing grout, but no grout bulb was observed beneath any of the 2.44 m long drilled shafts. In the case of group 2 shafts the grout again fl owed up alongside the shaft, thus increasing the shaft diameter during the first stage of grouting. During the second and third stage of grouting the grout again flowed up alongside the shaft and encircled the grout zone formed by the previous grouting. I n the case of group 2 east shaft (SG2 E) the second stage of grouting formed an accumulation only on one side above the tip of the shaft. The pressured grout always flowed along the least resistant path and accumulated at the loosest zone. The soil prepar ation in the test chamber or pulling of the casing after concrete placement may have created a relatively loose pocket on one side at the bottom of shaft SG2 E However, no such grout accumulation was observed at this location during the first stage of gro uting (Figure 5 9 ). Moreover, the grout accumulation was not attached to the side of the shaft, and hence its contribution to the capacity of the shaft was considered negligible. 5.2 Analysis of Experimental Post Grouted Drilled Shaft Group Behavior This S ect ion discusses the experimental behavior of tip grouted drilled shafts based on the load deformation data and soil stress measurement during different stages. Negligible ground surface deformation (<1mm) and soil stress change (Table 5
161 2) was observed near the cham ber wall during tip grouting and load tests which suggested that the chamber boundary effect was little, if any, in both group tests. From the group load tests prior to tip grouting the skin resistance of each shaft was found to be in the range of 13.3 16.5 kN (3 3.7 kips ) for group 1 shafts (2.44 m) and 4 1 46 kN (9.2 10.4 kips) for group 2 shafts (3.96 m) Figure s 5 10 and 5 1 1 displays the load versus vertical displacement response of group 1 and group 2 shaft s respectively during the group lo ad test s after tip grouting The maximum displacement of soil at the center of the group (group 1: 1 4 mm group 2: 9.6 mm ) was much less than the average displacement of shaft s (group1: 46 mm group1: 25 mm ), which suggests that the shafts behaved indepen dently during group loading In group 1, the load carr ying capacity of the s outh shaft was much higher than the other shafts because of the larger skin resistance due to the grout zone formed alongside the shaft (Figure 5 8 ). However, the load displacement responses of the shafts were nearly parallel at large displacements (above 25 mm) suggesting that the rate of tip mobilization was nearly the same for all the shafts. In case of group 2, the load carried by shafts were significantly different ; the west s haft carried the most load and the east shaft carried the least load at the same displacement, which was attributed to the difference in tip area (as observed in Section 5. 1 6 ) of the shaft after grouting As the maximum sustained grout pressure observed w as more or less the same for all shafts (Table 5 1) the effect of preloading (tip stress mobilization) should be similar. Figure 5 1 2 shows the load displacement response of south and west shafts in group 1 during the group and individual load tests Sim ilarly, shown in Figure 5 1 3 is the response of west and east shafts in group 2 under group and individual loading It is
162 e vident from the load displacement response s ( Figure s 5 12 and 5 13) that the shaft behavior w as similar in both scenarios. The respon se during individual loading was essentially the reloading of the response during group loading. D uring individual loading of shafts in each group, the displacements of surrounding shafts were insignificant as shown in Figures 5 1 4 and 5 1 5 which revea ls very little, if any, interaction betw een the shafts during loading. Note that if group inter action existed, loading on one shaft would cause notable displacement of other shafts in the group due to shear transfer through the soil. Figures 5 16 show s the vertical stress measurement beneath group foot print during the group load test s Note that the stress above 620 690 kPa (90 100 psi) could not be obtained due to the capacity of earth pressure cells being 690 kPa (100 psi). In both groups, the increase in vertical stress directly below the center of groups was significantly less er than the stress below the shafts unlike the jetted and grouted pile groups However, relatively larger stress increase at 0.61 m and 0.91 m below the center of group footprint ver sus the stress at 0.305 m ( below the center ) reveals that there was fairly small superposition of stress es transferred through the tip s of shafts (especially in group 2; where tip area increased by grouting). But th is stress overlap was much small in compa rison to jetted and grouted pile groups and hence the tip behavior of the groups may still be considered as individual (i.e., not block). In addition, the stress increase beneath the adjacent shafts during the individual loading of shafts in a group was li ttle, if any, as shown in Figure 5 1 7 The experimental data (load displacement responses, and observed stresses and displacements) revealed that the group failed through individual failure of each
163 shaft within the group. There was only small stress super position at the tip although some of the shafts had bigger grout bulbs at their tips (group 2). The minimal group interaction was attributed to little, if any, change in soil stresses and stiffness within the footprint area of the group as a resu lt of tip grouting. Hence, the resistance of the group is suggested to be equal to the single post grouted drilled shaft resistance the number of shafts in the group. Or, the axial group efficiency of the post grouted drilled shafts at typical 3D spacing is equal to one (1). Fig ure 5 1 8 presents the change in residual horizontal stress at the bottom side of shaft s caused by the construction axial loading, and tip grouting. In the cases of group 1 shafts the shaft construction and axial loading (prior to grouting ) caused a small increase in horizontal stress, but the stress change due to grouting was insignificant O n the other hand for group 2 shafts the shaft construction caused a decrease in horizontal stress, which may be because the casing for the shafts ( 3 .96 m shafts) had to be partially pulled during the concrete pouring (after 1.2 m or 4 ft of placement) to minimize the force required for the casing extraction, whereas for group 1 shafts ( short shafts), t he casing was pulled after the completion of the c oncrete placement. The partial pulling of the casing could have relieved the horizontal stress around the bottom of the shaft because the hydrostatic pressure exerted by fresh concrete at a relatively shallow depth (1.5 m) on the shaft wall was small. This horizontal stress reduction was regained after the first stage of grouting. The second stage of grouting caused an additional increase in horizontal stress (Fig ure 5 1 8 ). The increase in horizontal stress was also larger at 3.96 m than at 3.44 m, signifyi ng that the first stage of grouting may have prevented the upward flow of the grout during the
164 second stage of grouting. Although the horizontal stresses at the bottom sides of the shafts were different after construction in group 1 and group 2 shafts, the upward grout flow clearly occurred in both cases during the first stage of grouting. Similar upward grout flow was reported in full scale shafts by several investigators ( Duan and Kulhawy 2009; Muchard and Farouz 2009; Mullins and Winters 2004 ) and attrib uted to the lower horizontal stress around the shaft inherent in the installation process (drilling the shaft hole). Therefore, the grout pattern observed in the present study was not a result of the small scale effect and associated lower overburden stres ses. It is evident from the Fig ures 5 9 and 5 1 8 ( radial stress increase) that the grout zone also expanded laterally (in a somewhat cylindrical manner) in addition to flowing in the upward direction. However, it is clear that the tip grouting process was not represented by spherical cavity expansion beneath the shaft tip and the tip area increase was due to the accumulation of upward flowing grout around the shaft tips. The difference in the load displacement responses of the shafts as observed in Fig ure s 5 10 and 5 11 was attributed to the different final shaft tip areas. The experimental study also suggested that the shearing resistance of soil beneath the shaft tips was not necessarily improved (e.g., an increase in mean stress or relative density) thr ough the tip grouting process because no grout bulb developed and no grout soil mixture occurred beneath any of the shafts tested. Therefore the capacity of grouted shafts under permissible service displacements depended mainly on the pre stressing effect (i.e., the change in soil stiffness upon reloading) and increased tip area. T he experiments revealed that the stage grouting improved shaft tip resistance by preloading the shaft tip (i.e., the grouting process) and increasing the bottom side radius
165 of the shaft, however irregularly. The grout pressure measured during re grouting (stage 2 or 3 ) could not be used to estimate skin friction because the grout may not have covered the entire bottom area of the shaft, as observed in the experimental study.
166 Tab le 5 1 Drilled shaft s grouting data Group Length m Shaft ID Grout pressure, kPa a Grout volume m 3 (gal) Stage 1 Stage 2 Stage 3 Stage 1 Stage 2 Stage 3 S G1 2.44 S 448 0.064 (17) N 448 0.002 (0.5) W 448 0.002 (0.5) E 448 0.011 (3) S G2 3.96 S 1241 1724 0.023 (6) 0.023 (6) N 1034 1241 1241 0.023 (6) 0.023 (6) 0.011 (3) W 1103 1379 1241 0.023 (6) 0.023 (6) 0.011 (3) E 1172 1517 0.023 (6) 0.023 (6) a After deducting the pres sure lose in grout distribution system (138 kPa) from the measured pump pressure Table 5 2 Increase of horizontal soil stress at shaft tip elevation during grouting & group test Group Tip grouting Group load test 15 cm away from shaft 15 cm away from chamber wall 15 cm away from shaft 15 cm away from chamber wall SG1 32 kPa 2 kPa 93 kPa 10.5 kPa SG2 668 kPa 32 kPa 106 kPa 1.5 kPa
167 Figure 5 1. PVC casing positioned before filling the test chamber (Photo courtesy of author, Sudheesh Thiyyakkan di) Figure 5 2. Test chamber in fully filled state (Photo courtesy of author, Sudheesh Thiyyakkandi)
168 Figure 5 3. Reinforcing cage and grout distribution system (Photos courtesy of author, Sudheesh Thiyyakkandi)
169 Figure 5 4. Pulling the casing ou t (Photos courtesy of author, Sudheesh Thiyyakkandi) Figure 5 5 Group load test setup (Photo courtesy of author, Sudheesh Thiyyakkandi)
170 Figure 5 6 Tip grouting (Photo courtesy of author, Sudheesh Thiyyakkandi) Figure 5 7 Setup for individual shaft loading (Photo courtesy of author, Sudheesh Thiyyakkandi)
171 Figure 5 8. Excavated group 1 shafts (Photos courtesy of author, Sudheesh Thiyyakkandi)
172 Figure 5 9 Excavated group 2 shafts (Photo courtesy of author, Sudheesh Thiyyakkandi) Figu re 5 10. Load displacement response of group 1 shafts 0 20 40 60 80 100 120 140 0 5 10 15 20 25 30 35 40 45 50 55 60 Load (kN) Displacement (mm) North South West East
173 Figure 5 11. Load displacement response of group 2 shafts Figure 5 12. Load displacement response of group 1 shafts during the group and individual load tests. 0 50 100 150 200 250 300 0 5 10 15 20 25 30 35 Load (kN) Displacement (mm) North South West East 0 20 40 60 80 100 120 140 160 180 200 0 10 20 30 40 50 60 70 80 90 100 Load (kN) Displacement (mm) South during group test South during individual test West during group test West during individual test
174 Figure 5 13. Load displac ement response of group 2 shafts during the group and individual load tests Figure 5 14. Displacement s of all the shafts in group 1 during south shaft loading 0 40 80 120 160 200 240 280 320 360 0 10 20 30 40 50 60 Load (kN) Displacement (mm) East during group test East during individual test West during group test West during individual test 0 5 10 15 20 25 30 0 1 2 3 4 5 6 7 8 9 10 Displacement (mm) No of load steps S N E W
175 Figure 5 15. Displacements of all the shafts in group 2 during the individual shaft loa ding
176 Figure 5 16. Vertical stress measurement beneath group footprint during the group load tests
177 Figure 5 17. Vertical stress measurement beneath group footprint during the individual shaft loading
178 Figure 5 1 8 Variation of residual horizonta l stress after different stages measured using earth pressure cells at 15 cm (0.5 ft) away from shafts
179 CHAPTER 6 INDIVIDUAL RESPONSE OF POST GROUTED DRILLED SHAFTS This C hapter discusses the finite element modeling and the prediction approach developed for the a xial capacity of individual post grouted drilled shafts. The objective of the finite element analysis was to investigate the load transfer mechanism at the tip of grouted shafts. Experimental group study of p ost grouted drilled shafts at typical 3D spacing discussed in C hapter 5 revealed that the shafts behaved individually during axial group loading (i.e., little/no group interaction). Therefore, individual experimental shafts (both 2.44 m and 3.96 m long) w ere modeled in PLAXIS 2D and compared with the ex perimental results. The experimental observations the load transfer mechanism s captured from numerical analyses and the reported case histories were then used to develop a tip resistance prediction approach for the grouted shafts. 2 6.1 Numerical Modeling of Post Grouted Drilled Shafts Numerical modeling of the base grouted drilled shafts was performed using the two dimensional finite element software PLAXIS 2D The construction, first stage of grouting, and top down load testing of both the 0.216 m (8.5 in) diameter x 2.44 m (8 ft ) long and 0.216 m (8.5 in) diameter x 3.96 m (13 ft ) long shafts were modeled in PLAXIS. The second stage of grouting ( group 2 ) was not modeled because the grouting results (increase in tip radius and preloading effect) were irregul ar. Note that the tip radius increase ( Fig ures 5 8 and 5 9 ) after the first stage of grouting was nearly axisymmetric, and hence, the problem could be simplified as a two dimensional 2 Major portion of this Chapter is from the article: Thiyyakkandi, S., McVay, M., Bloomquist, D., Lai, P. (201 3 E xperimental study, numerical modeling of and axial prediction approach to base grouted drilled shafts in cohesionless soils Acta Geotechnica Spr inger D OI : 10.1007 /s11440 013 0246 3 With permission from Springer.
180 problem. An axisymmetric model using 15 node triangular elements was crea ted to simulate the drilled shaft and soil in the test chamber, as shown in Fig ure 6 1 Both the radial and vertical displacements were restricted along the periphery of the test chamber. The bottom mesh boundary was placed sufficiently deep under the shaf t tip such that its influence was negligible, to ensure that no stress change or displacement occurred near the bottom boundary during the different simulation stages, such as construction, grouting, and axial loading. Vertical movement was restricted but radial displacement (i.e., horizontal rollers) along the bottom boundary. 6.1.1 Material Models and Parameters The sand was modeled with the Hardening Soil (HS) constitutive model and the drilled shaft was modeled as a linear elastic material. As mentioned earl ier, the HS model considers three different moduli, E 50 from the primary deviatoric loading, E ur from the elastic unloading/reloading and E oed from the primary oedometer lo ading, to define soil stiffness. The model incorporates the stress level dependency of the stiffness parameters according to a power law controlled by a dependency parameter m E 50 ref = 3.46 x 10 4 kPa (secant modulus corresponds to 50% mobilization of the shear strength) obtained from the standard drained triaxial compression test was use d in the preliminary analysis. However, to match the FEM responses with the vertical soil stress measured by earth pressure cells at a depth of 0.305 m (1 ft) below the shaft tips during grouting and the axial top down load testing, E 50 ref had to be increa sed by 20%. A similar back calculation approach has been suggested by Youn and Tonon (2010) for the FEM modeling of post grouted drilled shafts. The materials parameters used for the sand and drilled shaft in the analysis are given in Table 6 1 The angle of internal friction used in the analysis was the peak friction angle obtained from direct shear tests at 50% relative
181 density and dilation angle, while was estimated using the Bolton (1986) model, ( c )/0.8 where c is the critical state frictio n angle and measures approximately 31 o for the test sand. The PLAXIS default settings of E ur ref = 3 E 50 ref and E oed ref = E 50 ref were used in the analysis. Although the cohesion of the silty sand used in the study was zero, a small cohesion value (0.345 kPa or 0.05 psi) was used in the analysis for the numerical stability of the model, as suggested in the PLAXIS manual. 6.1.2 Modeling of Construction, Tip Grouting and Top Down Load Tests In accordance with the experimental study, the FEM analysis began with the deactivating the soil elements representing the shaft. Subsequently, the concrete p lacement was modeled by activating the lateral pressure to the shaft wall, which was equal to the hydrostatic pressure induced by fluid concrete with a unit weight of 22.5 kN/m3 ( Youn and Tonon 2010) and the removal of the casing was simultaneously modele casing removal, and the shaft hole wall was free to deform under the lateral pressure (a resultant of the concrete and soil pressure) and alter the stress state of the soil around shaft, as in the actual construction process. The hydrated drilled shaft was then simulated by activating the linearly elastic concrete material at the shaft loc ation and deactivating the concrete fluid pressure. Interface elements were used at the interface between the shaft and soil to model the soil shaft interaction. The interface strength reduction factor R (the ratio of the shear strength of the interface el ement to the surrounding sand, or 0.85) was back calculated to match the initial skin resistance of the
182 shafts measured in the experimental study. Next, the tip grouting was simulated by applying positive incremental volumetric strains (i.e., expansion) to a linearly elastic zone [E = 689.5 kPa (100 psi) and = 0.45] at the shaft tip, as shown in Fig ure 6 1 This elastic zone simulated the cavity (voids) filled with grout during the commencement of grouting in the experimental study. The application of po sitive volumetric strains to the elastic zone, a PLAXIS recommended approach to the simulation of grouting (Ni and Cheng 2010; Thiyyakkandi et al. 2012 ) expanded the zone and exerted stress in all the directions. This expanding zone pushed the underlying soil downward (i.e., pre stressing) and the shaft upward. This upward pushing of the shaft caused the mobilization of skin resistance like that in the actual field grouting. The incremental application of positive volumetric strain was continued until an u pward shaft displacement of approximately 6.35 mm (0.25 in) was achieved, in accordance with the experimental study. In current industry practice, the applied grout pressure is either maintained or released during the hydration of the grout. However, it is impractical to maintain the full grout pressure because the stress relaxation in the soil beneath shaft causes a partial or full release of pressure. After the grouting simulation, three different grout pressure scenarios were considered in the analysis: 1) fully locked, 2) fully elastic grout zone by the grouting simulation (positive volumetric strain) was maintained. n the elastic zone was fully released, and in case 3, the stress was partially released. For cases 2 and 3, the grout pressure release was simulated by applying negative incremental volumetric strains (i.e., contraction) to the elastic grout zone until the vertical stress at shaft tip became zero or negligible
183 (case 2) or 50% of the applied grout pressure (case 3). The application of negative incremental volumetric strains contracted the elastic zone and relieved the stress developed by the grouting process Next, the elastic grout zone was replaced by linearly was used in the analysis because the grouting process is a large strain problem. After simulation of the grouting p rocess, the axial load test was modeled by applying incremental distributed loads on top of the shafts (with each increment measuring 344 kPa or 50 psi). For comparison un grouted shafts, both the 2.4 m and 3.96 m drilled shafts were also modeled in PLAXI S using the same material properties given in Table 6 1 The casting of the shafts was first simulated, and axial loading was then modeled by the activation of a distributed load on the shaft head, as described earlier. Shown in Fig ure 6 2 are the vertical stresses at 0.305 m (1 ft) below the shaft tip during the top down axial load test, obtained from the stress gages and the PLAXIS analysis of group 1 east shaft the 2.44 m long single stage tip grouted shaft. As mentioned earlier, the E 50 ref value obtain ed from the standard drained triaxial compression test had to be increased by 20% to match the numerical results with the experimental response. 6.1.3 Skin Resistance of Tip Grouted Shafts T he vertical stresses predicted by PLAXIS at the shaft tips were used to separate the side and tip resistance from the total applied load on the grouted shafts in the experimental study For instance, the experimental final skin resistance of the grouted shaft was estimated by subtracting the tip load (the unit PLAXIS tip resi stance multiplied by the measured final tip area) from the applied experimental top load, as shown in
184 Table 6 2. The final skin resistance of all the shafts was higher than the initial skin resistance (i.e., pre grouting), which was attributed to the incre ased surface area or radius of the shaft due to grout flowing up alongside the shaft and the increased horizontal stress on the bottom side of group 2 shafts (Fig ure 5 1 8 ). Table 6 3 presents a comparison of the grouted and un grouted skin resistances of t he drilled shafts from a number of full scale field tests in the United States. The skin resistance of grouted shafts was in the range of 85 % 113 % that of the un grouted shafts. It should be noted that the un grouted skin resistance for each case listed in Table 6 3 was the skin friction of the control shafts at the same site, and different skin resistance could therefore also be attributed to construction and spatial variability at the test sites. 6.1.4 Load Transfer Mechanism at Shaft Tip Fig ure 6 3 depicts the unit tip resistance v s. tip displacement from the numerical analysis of the un grouted and three grouted drilled shafts: 1) full locking of tip grout pressure, 2) full release of grout pressure, and 3) 50% release of grout pressure for the 3.96 m long ( group 2 ) drilled shafts. During the top down load testing, it was found that the unit end bearing of the base grouted shaft obtained for the three cases (grout pressure fully locked, partially released, and fully released) were nearly the same ( Figure 6 3 ) even though the load transfer mechanisms captured from the FEM analysis were different, as shown in Fig ure 6 4 In the case of locked in grout pressure, the mobilized end bearing was initially balanced by negative skin friction along the shaft ( Figure 6 4 case 1: initial). During the subsequent top down axial loading, the negative skin friction was first replaced by the applied load (case 1: stage 1). Further application of the load mobilized positive skin friction (acting upwards ; Figure 6 4 case 1: s tage 2), and additional end bearing was mobilized with additional loading (case 1: stage 3). In
185 the case of fully released tip resistance, no skin friction (i.e., negative skin friction) existed initially ( Figure 6 4 case 2 : initial). During axial loading positive skin resistance was mobilized, and the tip resistance was developed along a much stiffer path up to the initially applied grout pressure (reloading) under a small displacement ( Figure 6 4 case 2: stage 2), after which further tip resistance was mobilized at a much lower stiffness with additional loading (case 2: stage 3). The mechanism by which the load was transferred to the soil in the partially released case was similar to that of the fully locked case, as shown in Figure 6 4 ( case 3). The lo ad transfer mechanisms identified from the numerical analysis, as discussed above, were the same as the conceptual mechanisms of load transfer suggested by Mullins et al. (2006) Therefore, the present analysis validated the conceptual load transfer mechan isms and, as evident from Fig ure s 6 3 and 6 4 the net load capacity was essentially the same in the three scenarios Also shown in Fig ure 6 3 is the PLAXIS predicted tip response (thin solid line) for the un grouted shaft subject to top down axial load t esting. Note that if the tip was unloaded at 1200 kPa and subsequently reloaded, the resulting tip response followed a stiffer path up to 1200 kPa followed by a softer (less stiff) path. This behavior was the same as that of the un grouted shaft above the preloading pressure ( Figure 6 3 the long short dashed curve). This curve had the same stiffness response as the fully released tip grouted shaft curve ( Figure 6 3 dark solid line translated right ). As observed in the experimental study, neither grout bul b formation nor grout soil mixture occurred during the tip grouting process, indicating that tip grouting a shaft did not necessarily increase the principal stresses (e.g., cavity expansion) or strength characteristics (e.g., relative density or angle of i nternal friction) of the soil at the tip of
186 the shaft but only preloaded it. Consequently, the unit tip response of a base grouted shaft can be predicted if preloading stress (i.e., grout pressure) and the unit tip resistance vs. tip displacement curve for the un grouted shaft are known. However, the tip radius at the bottom of the shaft must be known to predict its resistance (or force) 6.2 Develop Axial Prediction Approach for Post Grouted Drilled Shafts 6.2.1 Estimation of Unit Tip Resistance vs. Tip Displacement Many investigators (De Beer 1984; Franke 1993; Lee and Salgado 1999) have suggested the use of cone penetration resistance ( q c ) in deep foundation design. In the case of granular soils, the drilled shaft unit tip resistance ( q b ) at tip displacements of 5 to 10 % has been related to cone tip resistance q c ( Lee and Salgado 1999 ) For instance, Eq uation 6 1 expresses the unit tip resistance at displacements of 10 % diameter ( q b0.1 ) in terms of q c and a coefficient Fig ure 6 5 gives the value of recommen ded by a number of researchers for bored pile/drilled shafts. (6 1) It is well known that the tip resistance displacement relationship is nearly hyperbolic in nature in deep foundations (Hirayama 1990; Fleming 1992) Chin (1970) has suggested that the asymptote of the hyperbolic tip resistance di splacement relationship can be represented in linear form by plotting the ratio of tip displacement to resistance against tip displacement. The inverse of the slope of the line represents the limit base resistance of the pile (CPT resistance q c ), which can be represented by: (6 2) In the above equation, represents the normalized base displacement (base displacement/shaft diameter) at any unit tip resistance q b and K b is the vertical intercept
187 on the /q b axis. Substituting = 0.1 (i.e., 10 % of the diameter) and Eq uation 6 1, into Eq uation 6 2, the intercept K b can be calculated as: (6 3) By substituting Eq uation 6 3 into Eq uation 6 2 the hyperbolic function representing the normalized unit tip resistanc e displacement relationship can be written as: (6 4 ) where the coefficient n is given by: (6 5 ) Fig ure 6 6 shows a comparison of the measured and predicted unit tip resistance vs. base displacement curves from Eq uation 6 4 for an un grouted shaft test performed at the Georgia Institute of Technology (Mayne and Harris 1993). The soil at the shaft tip is silty silica sand with 70 % sand and a clay fraction of less than 10 %. More details about the test site and l oad test can be found in the literature ( Mayne and Harris 1993 ; Lee and Salgado 1999 ) The relative density of the soil beneath the shaft tip is approximately 21 from Fig ure 6 5 (Lee and Salgado 1999). Reasonable agreement is observed between the predicted (using Eq uation 6 4 ) and measured response for the shaft. As discussed earlier, the shear strength of the soil below a grout tipped drilled (path AB, Figure 6 7 ) due to tip grouting. In other words, assuming full grout tip pressure
188 release along path BC ( Figure 6 7 ), subseque nt top down testing will exhibit two distinct tip stiffnesses, path CB ( q b g tip grout pressure) and path BD ( q b q g ) depending on the mobilized tip stress q b In the case of reloading, path CB may be characterized as (Timoshenko and Goodier 1970) or unit tip resistance q b vs reload displacement as: (6 6 ) w here k is stiffness, G 0 is the shear modulus, and modulus ( G 0 ) in Eq uation 6 6 can be represented by the mo dulus given by Randolph et al. (1994) for pile design ( Eq uation 4 1) In the case of path BD in Figure 6 7 the hyperbolic relationship proposed for a conventional drilled shaft, or Eq uation 6 4 can be used to assess tip stresses q b in terms of normalized tip displacements ( Figure 6 7 ): (6 7 ) where: (6 8 ) In Eq uation s 6 7 and 6 8 g ( Figure 6 7 ) is the prior permanent displacement of the tip grouted shaft due to the tip gr out pressure q g The first term in Eq uation 6 8 is the tip displacement from grouting, Eq uation 6 4 and the second term is the recovered elastic rebound, Eq uation 6 6 Again, Eq uation 6 6 is valid for tip pressures q b less than the planned tip grout press ures q g and Eq uation 6 7 is valid for tip pressures q b larger than prior tip grout pressures q g The tip capacity (load) at a given displacement can be
189 obtained by multiplying the unit tip resistance q b with tip area of the shaft. Figure 6 8 compares the predicted unit tip resistance dis placement curve for the 2.44 m long shaft (conventional and grouted) using Eq uations 6 4, 6 6 and 6 7 with the results from the finite element analysis. Similarly, Figure 6 9 shows a comparison for the 3.96 m long shaft. Be cause the cone penetration resistance ( q c ) at the shaft base was not measured, the relationship between q c and relative density ( D R ) suggested by Jamiolkowski et al. (2001) was used for the estimation. Details on all the parameters used for the prediction are listed in Table 6 4 It can be observed that the predicted unit end bearing displacement responses were comparable with the results from the numerical analyses. 6.2.2 Estimation of Tip Area Increase Due to Grouting As mentioned in C hapter 5 the experimental study showed that base grouting increased the tip radius of the shaft (grout flowed up alongside the shaft) in addition to preloading the soil beneath the shaft. Subsequently, an analysis of the tip area ratio A r (ratio of grouted to un grouted tip area) from the full scale case studies with measured CPT values ( q c ), volume of grout pumped, and maximum applied grout pressure was undertaken. Table 6 5 presents the details of the full scale field tests and relevant parameters used in the analysis. The first four full scale tests in Table 6 5 were from a site located in Clearwater, Florida. The soils at the site were loose to medium dense shelly sands. Cone penetration soundings were performed at the location of each shaft, and the CPT values ( q c ) at the elev ation of the shaft tips are given in Table 6 5 All of the test shafts were 0.61 m (2 ft) in diameter and 4.57 m (15 ft) in length. For the first two shafts (S1 FJ1 and S1 FJ2), grouting was performed using a flat jack type grout distribution system,
190 while a sleeve port type apparatus was used for the other two shafts (S1 SP1 and S1 SP2). The fifth and sixth test shafts in Table 6 5 were from another site (silty silica sand) in Clearwater, Florida. Both shafts had a diameter of 0.61 m (2 ft) and length of 4 .57 m (15 ft). A flat jack grout delivery system was used for one shaft (S2 FJ) and a Tube a Manchette (sleeve port system) was utilized in the other shaft (S2 TM). The CPT q c values at the shaft tips, measured grout pressures, and grout volumes for both s hafts are included in Table 6 5 All of the shafts in Clearwater test sites were constructed by the wet construction method, using polymer slurry for stabilization. The last full scale test (S5 S2) considered in the analysis was from Houston, Texas. The sh aft was 6.4 m long with a diameter of 1.22 m and was tipped in sand. Mineral slurry was used for the construction of shaft and grouting was carried out using a flat jack type grout distribution system. The Statnamic load testing method (Bermingham 2000) wa s utilized to estimate the axial response in all of the case studies considered here. More information about the site characterization and test programs for each field test can be found in the literature (Mullins et al. 2001 ; Mullins and Winter s 2004; Mull ins et al. 2006) For the analysis, the applied maximum sustained grout pressure ( q g ) was divided by the cone penetration resistance ( q c ) to create a dimensionless quantity termed the normalized grout pressure ( NGP ). Similarly, the volume of grout pumped ( V g ) was expressed as a non dimensional quantity [normalized grout volume ( NGV )] by dividing it by the third power of the shaft diameter ( D 3 ). Next, the measured mobilized tip stress p s from the load tests (Table 6 5 column 8), representing the transition from reloading to virgin tip resistance, was obtained from the change in slope in the log log plot of tip stress vs. displacement (Fig ure 6 10 ). For the calculation of tip stress p s the axial stress
191 (embedded strain gage data) at an elevation of 0.457 m 0.67 m (1.5 ft 2.2 ft) above the shaft tip was used instead of the shaft tip stress, which was assumed to be impacted by changes in tip area. Because the shaft tip was pre stressed by the applied grout pressure ( q g ), the transition from reloading to virg in tip resistance will occur at the stress of a magnitude equal to q g during subsequent axial loading. Note that if the tip area was not increased by the grouting process, the value of p s (from the embedded strain gage) measured during axial loading will b e equal to q g (recorded grout pressure) because the shaft tip area and cross sectional area at the strain gage location are the same. However, if the tip area was increased by grouting, the measured p s will be greater than q g (final shaft tip area > shaft cross sectional area at gage location). Accordingly, the increase in tip area, called the area ratio A r (i.e., the ratio of final to initial tip area) in Table 6 5 could be determined by dividing the measured mobilized tip stress p s by the reported tip gr out pressure q g Based on the analysis of seven full scale tests with known CPT values ( q c ) discussed above, it was found that the area ratio A r (Table 6 5 ) varied, even when the lengths and diameters of the shafts were similar. Moreover, it was found tha t the area ratio was correlated with the volume of grout pumped, maximum applied grout pressure, shaft diameter, and properties of the soil beneath the shaft tip. Several derived parameters in terms of normalized grout pressure ( NGP ) and normalized grout v olume ( NGV ) were considered in the study to check whether any correlation with the area ratio existed. It was found that the area ratio ( A r ) was most closely correlated to the ratio of normalized grout pressure to square root of normalized grout volume ( NG ), as shown in the scatter plot (Fig ure 6 11 ). In the ratio ( ), NGP accounts for the
192 grout pressure and soil properties ( CPT, q c is used in the normalization of grout pressure) and NGV for grout volume and shaft diameter D However, because t he tip area increase was due to the accumulation of upward flowing grout above the shaft tip, and not necessarily by the formation of a spherical bulb beneath the tip (discussed earlier), taking the square root may be an adequate modification to the term N GV (grout volume normalized by D 3 or sphere volume) to account for this actual scenario. Subsequently, the data were fitted with a hyperbolic relationship, Eq uation 6 9 with an R 2 of 0.85. The equation is applicable only to values less than 0.8 (based on available data). When the value is greater than 0.8, the area ratio ( A r ) should be taken as unity (conservative, or no tip area increase), as shown in Figure 6 11 For a specific field grout test (with a kno wn q c q g and grout volume ), A r is determined from Eq uation 6 9 and the estimated shaft tip resistance (stress) q b is found from either Eq uation 6 6 or 6 7 depending on the tip grout pressure q g and cone tip resistance q c while the estimated shaft tip r esistance (force) is found by multiplying q b by the design tip area and area ratio A r (6 9 ) 6.2.3 Comparison of Prediction Approach with Field Test Data Of interest was to compar e the proposed unit tip resistance ( q b ) displacement relationship (Eq uation s 6 6 and 6 7 ) with the results of the full scale ba se grouted drilled shaft tests. High quality full scale field data reported by Mullins et al. (2001), Mullin and along with estimated soil properties ( D r G 0 ), are given in Table 6 6
193 The first full scale fiel d test data considered for the comparison was fr om a site located at Clearwater ( Mullins et al. 2001). The soils at the site were loose to medium dense shelly sands with an average cone penetration resistance of 3200 kPa at the shaft tip. Fig ure 6 12 disp lays the predicted and measured tip load displacement response for the shaft. There are two predicted curves: one using the initial tip area of the shaft and the other an increased area (area ratio A r = 1.81, using Eq uation 6 9 for = 0.21). The pr edicted response considering the tip area increase matched the actual response well. The second test was conducted at the University of Houston in collaboration with University of South Florida on a 1.22 m diameter and 6.4 m long drilled shaft tipped in s and in the Houston region ( The pertinent geotechnical parameters for this case study are given in Table 6 6 The predicted tip load displacement curves using the initial tip area and increased tip area (area ratio A r = 1.76, using Eq uation 6 9 for = 0.23) are shown along with the measured response in Figure 6 13 which makes it clear that the predicted response with an increased tip area was in good agreement with the measured tip load displacement response. Finally, Dapp and Brown (2010) have reported a series of large diameter (2.29 m or 7.5 ft) base grouted drilled shaft tests for the Audubon Bridge project in Louisiana. All the shafts were approximately 61 m (200 ft) long and tipped in an alluvial sand and gravel forma tion with an average SPT blow count (N) of 64. The tip resistances of the grouted shafts were obtained by O cell load tests embedded near the shaft tip. Further details on the test program can be found in Dapp and Brown ( 2010) Figure 6 14 shows
194 the measur ed tip load displacement response for some of the shafts, and details on the parameters used in the prediction are given in Table 6 6 Because the diameter of the shafts was large, the possibility of enlarged shaft tip area due to grouting was negligible, which was also evident from the estimated area ratio ( A r ) value of nearly 1 ( Table 6 6 ). The predicted tip load displacement response ( A r = 1) is also shown in Figure 6 14 and is found to moderately match the measured responses.
195 Table 6 1 Material prop erties used in PLAXIS Parameter Sand Pile E (kPa) 2.48 x 10 7 E 50 ref (kPa) 4.16 x 10 4 E oed ref (kPa) 4.16 x 10 4 E ur ref (kPa) 1.248 x 10 5 P ref (kPa) 100 unsat (kN/m 3 ) 16 25 sat (kN/m 3 ) 19 Friction angle, 33 Dilation angle, 2 ur 0.25 0.15 Power, m 0. 5 Table 6 2 Skin resistance of shafts before and after grouting Shaft Initial skin, S i (kN) Measured final tip area, A f (m 2 ) Tip stress (FEM) @ 25 mm a q bg (kPa) Tip load, Q bg = q bg A f (kN) Top load @ 25 mm a Q tg (kN) Final skin S f = Q tg Q bg (kN) ( S f / S i ) x 100 (%) S G1 E 16 0.0486 1100 53.5 71 17.5 109% S G1 W 14 0.0366 1100 40.3 55 14.7 105% S G2 E 41 0.0534 1866 99.6 144 44.4 108% S G2 S 43 0.0843 1866 157.3 206 48.7 113% Note: Tip stress (column 4) from PLAXIS analysis a Top displacement
196 Table 6 3. Comparison of grouted and un grouted skin resistance from full scale tests in United States Site, shaft description Skin resistance (kN) (Grouted/ Un grouted)x100% Un grouted c Grouted Clearw ater, Florida, S1 FJ1 a 863 818 95% Clearwater, Florida, S1 FJ2 a 863 787 91% Clearwater, Florida, S1 SP1 a 863 774 90% Clearwater, Florida, S1 SP1 a 863 880 102% Clearwater, Florida, S2 FJ a 827 791 96% Clearwater, Florida, S2 TM a 827 827 100% Palm Beach, Florida, S3 LT3 a 28023 31315 112% West Palm Beach, Florida, S4 LT2 a 6449 6410 99% NGES, Alabama, TS1 b 2606 2322 89% NGES, Alabama, TS2 b 2606 2295 88% NGES, Alabama, TS3 b 2606 2349 90% Houston, Texas b 3051 2580 85% Broadway Bridge Viaduct Iowa 7860 8536 109% a After Mullin et al. (2001) b After Mullin et al. (2004) c Skin resistance of un grouted shaft at the same site
197 Table 6 4 Shafts and relevant parameters used for the prediction Test name Grout pressure, q g (kPa) q c a (kPa) D R (%) b G 0 c (kPa) FEM ( 2.44 m long shaft) 448 4850 a 49 0.16 17646 0.3 FEM ( 3.96 m long shaft) 1172 6530 a 49 0.16 23934 0.3 a using the equation proposed by Jamiolkowski et al. 2001  b from Fig. 8, according to Lee and Salgado 1999  c using Eq. 7. Table 6 5. Details of full scale field tests and relevant parameters Shaft a Shaft dia D (m) Grout press. q g (kPa) Grout vol. V g (m 3 ) CPT q c (kPa) NGP = q g /q c NGV = V g /D 3 p s (kPa) A r = p s /q g S1 FJ1, Clearwater, FL 0.61 586 0.05 2850 0.206 0.220 787 1.34 0.44 S1 FJ2, Clearwater, FL 0.61 462 0.107 3200 0.144 0.471 804 1.74 0.21 S1 SP1, Clearwater, FL 0.61 1138 0.165 2300 0.495 0.727 1266 1.11 0.58 S1 SP2, Clearwater, FL 0.61 1220 0.086 2400 0.508 0.379 1198 0.98 0.83 S2 FJ, Clearwa ter, FL 0.61 683 0.217 2100 0.325 0.956 1129 1.65 0.33 S2 TM, Clearwater, FL 0.61 862 0.18 2100 0.410 0.793 1403 1.63 0.46 S5 S2, Houston TX 1.22 1157 0.255 13400 0.086 0.140 2134 1.84 0.23 a After Mullins et al. 2001, 2004, and 2006
198 Table 6 6 Sha fts and relevant parameters used for the prediction Test name Shaft dia. D (m) Grout press. q g (kPa) Grout vol. V g (m 3 ) q c (MPa) D R (%) d G 0 e (MPa) Ar f Clearwater, Florida, S1 FJ2 0.61 462 0.107 3.2 33 b 0.2 29 0.3 1.81 Houston, Texas 1.22 1157 0.255 13.4 64 b 0.15 49.3 0.3 1.76 Audubon Bridge site, Louisiana 2.29 5200 0.23 1 24 a 50 c 0.16 203 0.3 1 1.06 a using the equation proposed by Jamiolkowski et al. ( 2001 ) b using the equation and chart suggested by Jamiolkowski et al. ( 2001 ) c using the SPT D R correlation suggested by Cubrinovski and Ishihara ( 1999 ) d from Fig ure 6 5 according to Lee and Salgado (1999) e using Eq uation 6 7. f using Eq uation 6 10
199 Figure 6 1. Typical Finite element discretization
200 Figure 6 2. Vertical stress at 0.305 m displacement during axial load test ( Group1 East shaft ) Figure 6 3. Unit tip resistance versus tip displacement for un grouted and grouted shafts from FEM 0 100 200 300 400 500 600 700 800 900 0 5 10 15 20 25 Vertical stress (kPa) Shaft top displacement (mm) tip gage stress (Group 1 East ) FEM 0 500 1000 1500 2000 2500 0 5 10 15 20 25 30 35 40 45 50 Unit tip resistance (kPa) Displacement (mm) Ungrouted shaft (loading) Grouted shaft (fully locked) Grouted shaft (50% released) Grouted shaft (fully released) Ungrouted shaft (reloading)
201 Figure 6 4. Mechanism of load transfer during axial loading of base grouted shaft
202 Figure 6 5 Values of recommended by various investigators and from load tests Figure 6 6. Predicted and measured response for Georgia Tech load test 0 1000 2000 3000 4000 5000 0 50 100 150 200 Unit tip resistance (kPa) Base displacement(mm) Measured Predicted Diameter = 0.76 m Length = 16.8 m CPT, qc = 6500 kPa Relative density = 25%
203 Figure 6 7 Conceptual normalized tip resistance displacement plot Figure 6 8. Comparison of unit tip resistance displacement response (2.44 m long shaft) 0 200 400 600 800 1000 1200 1400 1600 0 5 10 15 20 25 30 Unit tip resistance (kPa) Displacement (mm) FEM (grouted) FEM (ungrouted) Predicted (grouted) Predicted (ungrouted)
204 Figure 6 9. Comparison of unit tip resistance displacement response ( 3.96 m long shaft) Figure 6 1 0 Determination of mobilized tip stress, p s from log log plot of tip stress (embedment strain gage) vs dis placement (shaft: S1 FJ1 ) 0 500 1000 1500 2000 2500 3000 0 5 10 15 20 25 30 Unit tip resistance (kPa) Displacement (mm) FEM (grouted) FEM (ungrouted) Predicted (grouted) Predicted (ungrouted)
205 Figure 6 1 1 Area ratio ( A r ) versus NGP / V Figure 6 12. Predicted and measured tip load displacement response ( Clear w ater FL ) 0.00 0.50 1.00 1.50 2.00 2.50 3.00 0 0.2 0.4 0.6 0.8 1 A r 0 100 200 300 400 500 600 0 5 10 15 20 25 30 35 40 45 50 Tip load (kN) Displacement (mm) Measured (Mullins et al. 2001) Predicted (using initial tip area) Predicted (using increased tip area, Ar=1.81)
206 Figure 6 13. Predicted and measured tip load displacement response (Houston) Figure 6 1 4 Predicted and measured response (Audubon Bridge project, Louisi ana) 0 1000 2000 3000 4000 5000 6000 7000 0 20 40 60 80 100 120 Tip load (kN) Displacement (mm) Measured (Mullins and Winters 2004) Predicted (using initial tip area) Predicted (using increased tip area, Ar=1.76) 0 5 10 15 20 25 30 0 20 40 60 80 100 120 140 160 Tip load (MN) Displacement (mm) T2 Measured (Dapp and Brown 2010) T4 Measured 1W Measured 4W Measured 11E Measured Predicted (area ratio, Ar = 1)
207 CHAPTER 7 COMPARISON OF SIDE AND TIP GROUTED VERSUS TIP ONLY GROUTED FOUNDATIONS This C hapter presents a comparison of the effectiveness of side and tip grouting v s. tip only grouting on improving the axial resistance of deep foundations in individual and g roup placements. The results of the experimental and numerical investigations discussed in the C hapters 3 to 6 were used for the comparison. The difference in soil stress state in the vicinity of pile/shaft, the maximum tip grout pressure and grout bulb fo rmation, the axial response, and the group interaction in both scenarios are discussed in the following S ections. 7.1 Residual Horizontal Stress A round Deep Foundations The experimental and numerical analysis of jetted and grouted piles (i.e., side and tip gr outed deep foundation) suggested a significant increase in residual horizontal h ) and shear modulus of soil in close proximity of the pile s This is due to the fact that the side grouting of a pile with attached membranes resembles a cylindrical cavity expansion problem. Note that the membranes confine the grout zone and allow the radial expansion of the zone, which results the horizontal stress to become the major principal stress. Without the membrane, the grout will flow along the least resistance paths and hence the horizontal stress and shear modulus of the s urrounding soil are not necessarily improved. C onsequently, the piles undergone side grouting with membrane confinement have much higher ultimate side resistance compared to traditional deep foundations. In the case of cast in place foundations (e.g., dril led shafts) without side grouting, the horizontal stress is the minor principal stress. The installation process ( drilling the shaft hole ) diminishes the horizontal stress around the shaft As a result, such foundations mobilize the least skin resistance u nder axial loading.
208 7.2 Maximum Tip Grout Pressure and Grout Bulb Formation During tip grouting of a foundation, the grout pressure pushes the foundation upward concurrent with compressing the underneath soil. The upward grout force mobilizes the available ski n friction of the foundation. For a foundation with very high side resistance, the grout zone initiated at tip will expands as a nearly spherical bulb with continuation of grouting, provided that the grout does not find any weak path to flow. In such cases the maximum grout pressure expected will be the spherical cavity expansion limit pressure at that depth and further grout bulb expansion will occur under the constant expansion pressure (steady state expansion). Th us, in general the maximum expected tip g routed pressure for a foundation will be the minimum of the pressure required to mobilize the skin resistance and the spherical cavity expansion pressure at that depth. Experimental study of grout tipped drilled shafts (tip only grouted foundations) revea led that during grouting, the grout flows up along the shaft soil interface (weakest path) after filling the void space beneath the shaft tip The upward flow of the grout was attributed to the low horizontal stresses in soil above the shaft tip (inherent in installation process) compared to the vertical stresses at the shaft tip Conversely, no such upward grout flow was observed in the case of jetted and grouted piles (side grouted foundations) during tip grouting The increased horizontal stress (major p rincipal stress) subsequent to side grouting and the enlarged side grouted zones prevented the upward flow of grout. Since the upward flow ing grout may reduce (fluid grout carries no shear) grout pressure to estimate skin resistance (grout pressure x effective tip area) may underestimate the actual skin resistance of the shafts.
209 Table 7 1 presents a comparison between the measured tip grout pressures and spherical cavity expansion limit press ures for jetted and grouted piles and post grouted drilled shafts. In Table 7 1 the spherical cavity limit pressure at the tip of pile/shaft was predicted using Yu and Houlsby (1991) closed form solutions and Salgado and Randolph (2001) charts. It is evide nt from the Table 7 1 that the maximum tip grout pressures developed for jetted and grouted piles were in the range of the predicted spherical cavity limit pressure at corresponding depths. Th is indicates that the skin resistance of the piles subsequent to side grouting was sufficient to develop high tip grout pressure and thus cause the formation of a nearly spherical tip grout bulb by the cavity expansion process as shown in Figures 2 7, 4 10, and 4 1. Note that the higher grout pressures for the jetted a nd grouted piles were only available after the higher horizontal stresses (i.e., major principal stress) were mobilized due to side grouting of the piles with the membranes. Whereas the maximum tip grout pressures measured for the drilled shafts (not side grouted) were much less than the spherical cavity expansion pressures (Table 7 1 ) For example, the tip grout pressures observed for 2.44 m long jetted and grouted piles were in the range of 1 517 17 93 kPa. In contrast, the grout pressures observed during tip grouting of 2.44 long drilled shafts were only one third (448 kPa or 65 psi) of that for the jetted and grout piles because of the lesser skin resistance This small er grout pressures w ere not sufficient to cause the spherical cavity expansion and henc e no such spherical bulb was formed beneath any of the tested post grouted drilled shafts (Figures 5 8 and 5 9). Since the spherical grout bulb formations (by cavity expansion process) will improve the mean stress and relative density of underlying soil, t he unit tip resistance of
210 jetted and grouted pile s will be significantly improved by tip grouting. But in post grouted drilled shafts neither grout bulb formation nor grout soil mixing (i.e., grout permeation) occurred below any of the test shafts. But th e tip area of the shafts were increased (especially in group 2 shafts) due to the accumulation of upward flowing grout a bove the shaft tips I n addition to increasing the tip area, the tip grouting in such foundations (without side grouting) only compresse s the soil immediately below the shaft tip and increase s its stiffness but not necessarily the soi shear strength. 7.3 Axial Resistance As discussed earlier, side grouting of a pile with membrane improves the ultimate side resistance of the pile and assis ts in developing greater tip grout pressure and spherical tip bulb formation during tip grouting. Spherical cavity expansion beneath shafts improves the stress state of the soil and increases tip area. This results in the mobilization very high tip resista nce during the subsequent axial loading. Thus the side and subsequent tip grouting of a foundation significantly improves its axial resistance. On the other hand, in tip only grouted foundations (post grouted drilled shafts) no grout bulb was formed benea improved because of the low grout pressure. Therefore, it is postulated that grouting of a drilled shaft tip just preloads the soil, and the grout pressure times the shaft tip area will be mobi lized under small deformations during subsequent axial loading (i.e., higher stiffness); however, the ultimate unit tip resistance at large displacement may not be altered. 7.4 Group Interaction In the case of side and tip grouted piles installed in a group at typical 3D spacing the experimental data suggested that grouting of adjacent piles within the group
211 increased the confining stress and shear modulus of the soil mass within the group, resulting in a very low shear strain development in the soil mass with in the group footp rint area under axial loading (Figure 7 1). However, outside the footprint, the horizontal stresses and shear modulus diminished quadratica lly with radial distance Consequently the group failed as block with uniform displacement within t he group and quadratically decreasing displacements outside the group ( Figures 4 16 and 4 17) Due to little confinement, a much higher shear strain pattern developed in the soil mass outside the footprint where the shear modulus greatly diminished due to reduction in horizontal stress (no longer principal st ress due to rotation of pole). Since the group behaves as a block during axial loading, the group efficiency factor for side shear will always be less than one (1) due to the reduced surface area of the group compared to the sum of the individual pile surface areas In contrast, the group foot print area is greater than the sum of individual pile tip areas, which results in group efficiency factor of greater than one (1) for the tip resistance .Consequen tly the group efficiency factor for total axial resistance (side + tip) depends on the dominancy of side or tip resistance. However it is suggested to estimate the axial group resistance using the approach discussed in the Section 4 3 instead of estimating the individual axial resistance and multiplying with group efficiency factor. Specifically, use the surface area of the block for estimating total side resistance and the effective block footprint area for estimating tip resistance T he experimental data on tip grouted drilled shaft groups suggested that there was little increase in radial stress and soil displacement around shaft tips during tip grouting which was attributed to the low grout pressures developed due to lower skin
212 resistance of shafts. The negligible increase in confining stress and soil displacement was not enough to improve the soil stiffness between shafts, and consequently, high shear strain developed in soil around the shaft t ip during axial group loading. Thus the groups failed throug h the individual failure of each shaft within the group (Figure 7 1) Hence, the resistance of the group is suggested to be equal to the single tip grouted shaft resistance number of shafts in the group
213 Table 7 1. Comparison between the measured tip grout pressures and spherical cavity expansion limit pressures Foundation type Pile/shaft ID Length (m) Width/ Diameter (m) Measured tip grout pressure (kPa) S pherical cavity expansion pressure (kPa) Yu and Houl sby (1991) Salgado and Randolph (2001) Jetted and g routed piles PG1 piles 2.44 0.203 1517 1793 1450 1620 PG2 piles 2.44 0.114 1689 2138 1450 1620 Field test piles 5.5 0.71 2000 2206 219 0 2 900 Post grouted drilled shafts SG1 shafts 2.44 0.21 6 448 1450 1620 SG2 shafts 3.96 0.216 1034 1241 2068 2900
214 Figure 7 1. Group behavior of jetted and grouted piles and post grouted drilled shafts Quadratic diminishing of horizontal stress High confining stress and shear modulus Low shear strain during axial loading High shear strain during axial loading Low shear modulus a) Jetted and grouted pile group Individual shaft failure Block failure b) Post grouted drilled shaft group
215 CHAPTER 8 CONCLUSIONS This dissertation encompass es the behavior of grouted deep foundations in individual and group placement. Focus has been made to characterize the axial response of jetted and grouted piles (i.e., side and tip grouted foundations) and post grouted drilled shafts (i.e., tip only grouted foundations) in cohesionless soils. Experimental and numerical studies were performed on both foundations and the following specific conclusions are dawn from the findings of the investigations. In case of jetted and grouted piles, the c ylindrical cavity expansion pressures (Yu and Ho ulsby 1991, Salgado and Randolph 2001) were found to be reasonable predictors of the expected grout pressure during side grouting. Similarly, the expected grout pressure during tip grouting was found to be equal to the minimum of the spherical cavity expan sion pressure or the pressure required for mobilizing the full side resistance of pile. A numerical analysis (PLAXIS 2D) of side and tip grouting as well as axial loading of the jetted and grouted precast piles showed that grouting developed the major prin cipal stresses in the radial direction and minor stresses in the vertical direction However, under vertical top down loading, the radial stress was found to diminish, but the minor principal stress around pile, which is equal to the vertical stress after grouting ( vg ) changed only slightly. Based on the experimental/numerical results for the vertical stresses after grouting, the ultimate unit skin friction along the pile may be predicted by the Mohr Coulomb strength for cohesionless soils. For the latte r, the magnitude of the unit skin friction is a function of depth, strength and relative density of the soil. Using the estimate of the ultimate skin and tip resistance, the axial load displacement responses of the piles were also predicted Because this w ork suggests
216 methodologies for the design of jetted and grouted precast piles, more testing is required to verify the results before implementation is recommended. Specifically, more studies on the stresses and mobilized resistance of the piles are warrant ed. T he constructability of the pile in typical Florida field conditions was verified by performing full scale field installation of two jetted and grouted piles. The noise and ground surface vibration measurements during the jetting and grouting operation s suggested that the pile is well suited for urban environment where the noise and vibrations during the construction operations are critical concern s Axial top down testing a jetted and grouted pile and a similar sized drilled shaft in the same field con dition showed that the unit skin friction for the pile is much greater than (2.6 times) that of drilled shaft. Similarly, combined torsion and lateral load testing of the pile using a full scale Mast arm structure identified that the new pile has very high torsional resistance and hence appropriate for Mast arm assemblies supporting highway signs and signals. However further studies to identify any reduction in unit skin friction associated with long term creep (e.g., one year), resulting in the redistribu tion of stress field around pile, are needed. Nevertheless, jetted and grouted precast piles are found to be a promising deep foundation system for the future. The group interaction of jetted and grouted piles at typical 3x pile width spacing was studied by performing two small scale group tests in cohesionless soils Measured load displace ment response of the piles under group loading revealed that the displacements of all piles were relatively uniform irrespective of the load carried by each pile Simila rly, the soil deformation at the center of the group was almost identical to the average displacement of the piles In addition, the vertical stress beneath the center of
217 the group was higher than the vertical stress increase recorded directly beneath pile s due to overlapping stress bulbs from individual piles All of these observations suggested that the piles behaved as a single block during axial loading The axial load displacement responses of the groups were also predicted using the same methodology s uggested for single piles. However further studies are required to investigate the group interaction at other spacing such as 4 D and 5 D ( D = pile width/diameter) In the case of post grouted drilled shafts, two small scale post grouted drilled shaft groups at 3D spacing were tested to study the group interaction behavior of post grouted drilled shafts The displacement of soil at the center of the group measured during group loading was much smaller than the average displacement of the top of the shaft. Mor eover, the vertical soil stresses measured beneath the shaft group during group loading showed little, if any, stress increase at the center of group versus directly under a shaft, unlike the jetted and grouted pile groups The axial response of the shafts in a group during the group and individual loading were essentially the same. Specifically, t he response during individual loading was the reloading of the response during group loading During individual loading of shafts in each group, the displacements of surrounding shafts were negligible. Based on the measured displacements and stresses during group and individual loading, it was suggested that the grout tipped drilled shafts behaved individually with little, if any, influence on another (i.e., group efficiency factor is 1). Since the study was limited to small scale group tests, full scale group tests need to be performed in future to validate the observed behavior. In addition, the experimental study of post grouted drilled shaft groups revealed tha t the grout flowed predominantly up along the shaft soil interface during base
218 grouting. Grout was also shown to bond to the shaft after hydration and cause an increase in tip radius and tip force. Although multiple stage grouting was a viable solution to preventing upward grout flow, the grout did not cover the entire tip area resistance. In addition, grout bulbs were not formed by spherical cavity expansion. Instead, t he grout flowed upward along the path of least resistance. Tip area (or radius) increase was due to the accumulation of upward flowing grout just above the shaft tip. No grout permeation or expansion into the soil beneath the shaft was observed in any of t It was concluded that the capacity of grouted shafts under permissible service displacements depended mainly on the pre stressing effect (i.e., the change in soil stiffness upon reloadi ng) and increased tip area. Finite element modeling of the grouted and un grouted shafts using PLAXIS 2D showed that the unit tip resistance displacement responses of the grouted shafts were the same as the reloading responses of similar sized un grouted drilled shafts initially loaded to the anticipated grout pressure. Numerical analysis also revealed that the tip resistance of base grouted drilled shafts with grout pressure either locked in or released were virtually the same. A simple approach to predic ting the tip resistance of base grouted shafts in cohesionless soils was suggested based on the results of this study. The approach utilizes cone penetration resistance ( q c ) and the small strain shear modulus ( G 0 ) of soil beneath the shaft, the shaft diame ter ( D ), anticipated grout pressure, and anticipated grout volume to predict both the final tip area and the unit tip resistance displacement response of a grouted shaft. The new method was applied to a
219 number of full scale case studies reported in the lit erature for comparison, yielding reasonable agreement. However, further validation of the proposed method (e.g., on site load testing) is required before it can be implemented in practice. A comparison of the behavior of jetted and grouted piles (i.e., si de and tip grouted foundations) versus post grouted drilled shafts (i.e., side only grouted foundations) showed that side grouting of a foundation prior to tip grouting has a significant role in improving the capacity. In addition to improving the stress s tate around pile and thus increasing the skin resistance, the side grouting also assists in developing a high grout pressure during tip grouting and preventing the grouting flow along the pile/shaft soil interface. Th is allows the grout zone to expand as a spherical bulb (i.e., cavity expansion) during tip grouting and improves the mean stress of the underlying soil, which consecutively improves the unit tip resistance of the pile or shaft. In case of group placement, side and tip grouting of adjacent piles within the group increases the confining stress and relative density of the soil mass within the group Therefore the soil mass within the group footprint area experience s very low shear strain during subsequent axial loading and the group fails as a sing le block with uniform group displacement But in case of tip only grouted foundations ( tip grouted drilled shafts ) negligible increase in radial stress and radial soil displacement occurs around shaft tips during tip grouting due to low tip grout pressur es A s a result high shear strain develops in soil around the shaft tip during axial group loading and the individual failure of each shaft in the group occurs. Thus the study proves that the side grouting of a foundation prior to tip grouting has a major role in improving the axial response in individual and group placement and the proposed hypothesis is validated
220 LIST OF REFERENCES AASHTO (198 2 ). Standard recommended practice for evaluation of transportation related earth borne vibrations, AASHTO Designa tion: R8 81 AASHTO (2010). LRFD b ridge d esign s pecifications, 5 th Edition, AASHTO. ACI Standard ACI 318 0 8 (200 8 ). Building c ode r equirements for s tructural c oncrete, American Concrete Institute, Farmington Hills, Michigan. Anderson, B. J., and Towns c onstitutive m odels for Florida s ands by d rained t riaxial t ests, C onstitutive M odels, GeoFrontiers 2005, GSP 139, ASCE, 1 12. ASTM D 1143/D 1143M (2007). Standard t est m ethods for d eep f oundations u nder s tatic a xial c ompression l oad, ASTM. ASTM D 2937 (2010). Standard t est m ethod for d ensity of s oil in p lace by the d rive c ylinder m ethod, ASTM e lement a nalysis in e ngineering a nalysis Prentice Hall, New Jersey Bermingham, f irst t en y Second International Statnamic Seminar, Tokyo, Japan, October 28 30, 1998, 457 464. ceedings of Physics Society, 57, 147 159. g routing p reloading of l arge p iles on s and In: Proceedings of 8th ICSME, Moscow, 20. s trength and d ilatancy of s ands Geotechnique 36(1), 65 78. c onstruction of c ontinuous f light a uger p iles, No.8, FHWA. p erformance of l arge d iam eter p iles by g routing Parts 1 and 2, Ground Engineering, 19(4), 9 15 and 19(5), 1121 1126. g routing t erminology Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 131(12), 1534 1542.
221 Bullock, P. J., Sch s hear s etup. I: t est p iles d riven in Florida Geoenvironmental Engineering, ASCE, 131(3), 292 300. e xpansion in c ohesive f rictional s o ils 358. d isplacement of g round c aused by s oil c ement c olumn i nstallation Engineering, ASCE, 1 31(5), 623 632. failure 92. v alue and relative density for sandy soils Soils and Foundations 39(5), 61 71. base post grouting piles of the Suramadu bridge Paper ID GTJ102 926. Audubon bridge modeling & design, ASCE, 1553 1562. rience with base grouted drilled shafts in the southeastern United State In: Proceedings of 10th international conference on piling and deep foundations, Deep foundations institute, Amsterdam 2006, article #1385, publication #77 (IC 2006). De Beer, E. ( conf. on soil mech. and found. Eng., G. Petrasovits, ed., Budapest, 307 318. axially loaded p iles in Europe Piles, ISSMGE Subcommittee. In: Proc. 13th European conf. on soil mech. and geotech. Eng., Prague, Czech Republic, 663 715. grouting of slurry drilled shafts in soil: Chinese experiences Equipment Expo, Contemporary Topics in Deep Foundations (GSP 185), Orlando, FL, ASCE; 47 54. project, School of Civil & Resource Engineering, University of Western Australia, Australia.
222 t hick c ylinder of s and: A l aboratory s imulation of the Pressuremeter t 424. F DOT (2002 a ). Section 455 St ructures Foundations, F lorida Department of Transportation Standard FDOT (2002b) Structures d esign gui delines for l oad and resistance f actor design Florida Department of Transportation Standard FDOT (2011). Structures Manual, 9, Florida Department of Transportation Standard. HEP 05 054 analysis 42(3), 411 425. of bored piles, including negative skin friction and horizontal loading I nt. G eotech. Seminar on D eep F ound. on B ored and A uger P iles, Van Impe, ed., Balkema, Rotterdam, Netherlands, 43 57. Gabr, M. A., Borden, R. H., Denton, R. L., and j etting i nduced d isturbance z one and a ssociated e cological i FHWA/NC/2006 09, North Carolina State University in Corporation with North Carolina Department of Transportation and Institute for Tra nsportation Research and Education. displacement of drilled shafts in sands In: A. Yeung and T. Felio, eds., Vol. 2, Geotech. Engrg. Div. ASCE, Reston, VA, 1039 1057. 618. n ew f oundation t echnique u sing p iles s ea led by c oncrete u nder h igh p Technical Conference. Houston, Texas, Vol. 2, Paper No. OTC 2310, 641 656. e ffect of p ile j ett ing and p Department of Transportation, in Corporation with US Department of Transportation and Federal Highway Administration. Hill, R. (1950). The mathematical theory of plasticity, Oxford University press.
223 H settlement analysis for bored piles using hyperbolic transfer functions 64. Hughes, J. M. O., Wrath, C. P. and Windle, D. (1977). Pressuremeter G e otechnique 27( 4 ) 455 477. J s haft f oundation d efects: i dentification, i maging, and c CFL/TD 05 007. r elative d ensity and s hear s trength of s ands from Cone p enetration t est and Flat d ilatometer t Proceedings of the Symposium, Reston, VA, ASCE, 201 238. Joer, H. A., Randolph, M. F., and Gunasena, U. (1998 m odeling of the s haft c apacity of g routed d riven p 21(3), 159 168. 4(1), 1 23. ty p roblem of l arge d iameter b ored p Journal of Geotechnical and Geoenvironmental Engineering ASCE, 135(2), 237 245. a xially l oaded p ipe p iles in s and Geotechnical Engineering, ASCE, 117(2), 272 296. shafts in sand Journal of geotechnical engineering, 105(GT1), 31 47. Lai, P., McVay, M., Bloomquis i nnovative p refabricated p ile i nstallation m ethod u tilizing j etting and p ressure g routing GeoFlorida 2010: Advances in Analysis, Modeling & Design (GSP 199) 2010, ASCE. 1592 1601. Lee, J. H., and Salgad Journal of Geotechnical and Geoenvironmental Engineering, 125(8), 673 683. c Hall, New Jersey. Mayne, P. displacement behavior of drilled shaft foundation in Piedmont residium 30 2175, Federal Highway Administration, Washington, D.C.
224 Pre stress ed c oncrete p ile i nstallation u tiliz ing j etting and p ressure g routing Final Report submitted Florida Department of Transportation, Fl orida McVay, M. C., O'Brien, M., Townsend, F. C., Bloomquist, D. G., and Caliendo, J. A. (1989). "Numerical a nalysis of v ertically l oaded p ile g roups," Proceedings of Foundation Engineering Congress, ASCE, Northwestern University, Illinois, 675 690. a pparatus for m easuring the s trength of s oils in p Thesis, University of Illinoi s, Illinois, USA. soils Journal of S oil M echanics and F oundation D ivision, 85(SM6), 1 29. c apacity and s ettlement of p ile f ournal of Geotechnical Engineering Division, ASCE, 102(3), 197 228. t est s haft p ost g routing and l oad t esting County, Florida f or post grouted shafts In: 34th annual Conference on deep foundations, Deep foundations institute, Kansas city. shaft tips phase I a Department of Transportation, Florida. a full scale load test program grouting drilled shaft tips phase II Final report submitted to Florida Department of Transportation, Florida. post grouted drilled shaft in cohesionless soils rnal of Geotechnical and Geoenvironmental Engineering, 132(4), 478 87. NCHRP (2003). S tructural supports for Highway signs, Luminaires and Traffic signals NCHRP report 494, National Cooperative Highway Research Program. c ap acity of a uger cast p iles in s Journal of Geotechnical Engineering ASCE, 117(2), 331 345.
225 Ni, J. C m odeling g rout e fficiency of l ifting s tructure in s oft c lay 229. a ugured p ile p ractice o Transportation Research Record No. 1447, Design and Construction of Augur Cast Piles and Other Found ation Issues, Transportation Research Board, National Research Council, Washington, D.C., 3 9. s hafts: Construction p rocedures and d esign m IF 95 025 Palmer, A. C. (1972 Geotechnique, 22(3), 451 457. s haft g routed p iles and b arrettes ings of 14th Southeast Asian Geotechnical Conference, Hong Kong, Vol.1, ISBN 90 5809 250 X, 407 412. d eformation of v ertically l oaded p ), 1465 1488. d riven p iles in s and Geotechnique, 44(3), 427 448. model friction piles ical Journal, 1(2), 81 93. load transfer curves Puerto Rico. nds based on CPT results Am. conf. of soil mech. and found. eng, Guadalajara, 3,1261 1274. eoenvironmental Engineering, 123(4), 344 354. c hamber s ize e ffects on p enetration r esistance in s Journal of Geotechni cal and Geoenvironmental Engineering 124(9), 878 888. c avity e xpansion p ressure and p enetration r esistance in s 251 265.
226 Salgado, R., and Randolph, International Journal of Ge omechanics, ASCE, 1(2), 175 192. h ardening s oil m odel: f ormulation and v erification tor, Beyond 2000 in Computational Geotechnics, Balkema, Rotterdam, 280 290. a ssessment of c hamber s ize e ffects in the c alibration of i n s itu t ests in s 445. round v ibrations c aused by p ile i 4th International Conference on Piling and Deep Foundations, Italy, 497 502 Shestopal, A. O. (1959). Jetting of p ipes, p iles, and s heet p iles Hydro project Institute, Moscow, USSR. Smi th, C. B., and Bloomquist, D. (2010) Horizontal Permeameter : t he design, development and testing of a horizontal permeability laboratory test, Magna Cum Laude Honors Research Project University of Florida. grouti ng on the load bearing capacity of bored piles foundation engineering, A. A. Balkema, Helsinki: Rotterdam, Finland, 167 70. s oil m ovements from p ile d Periodical on Structural Design and Construction, ASCE, 11(2), 80 85. escavadas submetidas a sucessivas provas de carga, In: Proc., 10th Brazilian conf. on soil mech. and found. eng., Foz DoIguacu, 1, 3 9. p redicted r esponse of a n ew j etted and g routed p recast p ile with m embranes in c ohesionless s oils l and Geoenvironmental Engineering, doi: 10.1061/(ASCE)GT.1943 5606.0000860. numerical modeling of and axial prediction approach to base grouted drilled shafts in cohesionle 013 0246 3. Timoshenko, S. P., and Goodier, J. N. (1970). Theory of Elasticity, McGraw Hill, New York.
227 114(3), 32 6 336. Mech anics and Found ations Div ision 98 ( 3 ) 265 290. i n p iling: g round v ibration and n oise during p ile i International Deep Foundations Congress, Orlando, USA, ASCE Special Publication 116, 363 371. White, D. J., Sidhu, H. K., Finlay, T. C. R., Bolton, M. D., and Nagayama, T. (2000). i n p iling: t he i nfluen ce of p lugging on d International Conference of the Deep Foundations Institute, New York, 299 310 e ffects of p ile i nstallations on a djacent s NCHRP Synthesis 253, Transportation Research Board, National Research Council, Washington D.C. grouted drilled shafts: a case study at the Brazo River Bridge, TX 456 465. Yu, H. S. (2000). Cavity e xpansion m ethods in geomechanics, Klu wer Academic Publishers, London c avity e xpansion in d ilatant s oils l oading a 183.
228 BIOGRAPHICAL SKETCH Sudheesh Thiyyakkandi was born in Kerala state, India and remained in Kerala until he moved to United States to join the PhD program He earned his Bachelor of Technology (B. Tech) degree in Civil Engineering from University of Calicut Kerala, India in 2002. He received his Master of Technology (M. Tech) degree in Civil Engineering (Geotechnical) in 2004 from the University of Kerala, India. Soon after M. Tech, he joined Ideal Educational Society (IES) College of Engineering, Thrissur, Kerala, as Lecturer in Civil Engineering. In Augus t 200 5 he joined National Institute of Technology Calicut, Kerala, India as Lecturer/Assistant Professor. He was accepted by the Department of Civil and Coastal Engineering Department at the University of Florida for doctoral program in August 2008. He wor ked as graduate research assistant under Dr. Michael C. McVay while doing PhD. He completed his Doc tor of Philosophy in May 2013.