TRINITY COLLEGE DUBLIN
The Influence of Lubricant Slurries on Skin Friction Resistance in Pipe Jacking
A thesis presented to the University of Dublin, Trinity College for the degree of Doctor of Philosophy
Ciaran C. Reilly
School of Engineering Supervisor: Assoc. Prof. Trevor L. L. Orr 30th May 2014
Declaration I declare that this thesis entitled “The Influence of Lubricant Slurries on Skin Friction Resistance in Pipe Jacking” is entirely my own work and has not been submitted as an exercise for a degree at this or any other University. I agree to deposit this thesis in the University’s open access institutional repository or allow the Library to do so on my behalf, subject to Irish Copyright Legislation and Trinity College Library conditions of use and acknowledgement. Signed, _______________________________ Ciaran C. Reilly
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“Think positive! Together we are building the future.” -‐ Dr.-‐Ing. E.h. Martin Herrenknecht Chairman of the Board of Management, Herrenknecht AG
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Abstract The use of bentonite and polymer slurries as lubricants to reduce jacking force has long been established practice in pipe jacking and microtunnelling, with reductions in the skin frictional resistance component of jacking force of up to 90% being reported. The objective of this research is to improve the understanding of the mechanisms of action of these pipe-‐ jacking lubricants. Field case histories are presented where the influence of lubricants is shown. Laboratory studies are described where the effects of pressurised and unpressurised lubricant slurries in the interface between concrete jacking pipes and coarse-‐ grained and fine-‐grained soils are quantified. A novel triaxial testing apparatus is presented, which allows for the injection of lubricants onto the interface between rough concrete and a soil that is shearing at the constant volume interface friction angle. It is shown from the results of laboratory testing that the main beneficial effect of pipe jacking lubricants is the reduction of the radial effective stress acting on the pipeline through the generation and retention of excess pore water pressure in the soil near the interface. Results are compared to field case history data and conclusions are drawn. Additionally, the jacking forces observed during two recent microtunnelling projects in rock in Ireland are presented and analysed.
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Acknowledgements I had the privilege to meet and work with a great many remarkable people during my four years in Trinity College Dublin, and I wish to thank each and every one. In particular, I would like to thank my supervisor, Dr. Trevor Orr, for his constant support and encouragement during my time at TCD. His patience, advice and continuous effort were vital to the preparation of this thesis. The tireless assistance and helpful advice received from Martin Carney and Eoin Dunne in the TCD Geotechnical Engineering Laboratory, and all the technical staff in TCD, is gratefully acknowledged. I wish to thank all my postgraduate colleagues in the Department of Civil, Structural and Environmental Engineering, whom I had the pleasure of sharing offices and labs with, for friendship and support. I would like to thank colleagues from industry and academia who have assisted greatly with the practical aspects of this work, and who have been generous with their time: Barry McAllister and Robert McGoran of Terra Solutions Ltd., Ken McIntyre, Eamonn Sweetman and Gabrielle Cronin from J. B. Barry & Partners Ltd., Mark Brooks of Mudtech Ltd., Martin Sommes of Herrenknecht Internation LTD. (UK), Paul Wilkinson of Iseki Microtunnelling and Paul Willis of F&B Trenchless Solutions Ltd. Thanks are particularly due to Prof. Stephan Jefferis, Prof. George Milligan, JP Smith and John F. Nally, who assisted greatly in a personal capacity. The financial support received from the Irish Research Council is gratefully acknowledged, the work reported in this thesis would not have been possible without it. Finally, I dedicate this thesis to my parents, Chris and Catherine, for encouraging and facilitating my education over the years, and to Niamh, for endless support and patience.
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List of symbols
a
half contact width, b/2
AC
corrected area
b
contact width for pipe resting on base of stable bore
c'
drained cohesion of soil
cd
apparent cohesion from undrained tests
cu
undrained shear strength of soil
cur
residual undrained shear strength of soil
CU
coefficient of uniformity
CC
coefficient of curvature
C
cover depth = z0-‐DP/2
DP
product pipe diameter
DS
shield or microtunnelling machine diameter
Dr
relative density of soil
emax
maximum void ratio of a soil under consideration
emin
minimum void ratio of a soil under consideration
Ep
elastic modulus of product pipes
Es
elastic modulus of soil
F
axial force
fs
sleeve friction resistance in the cone penetration test
IL
liquidity index
L
length of jacked pipeline
M
skin resistance friction factor
N
tunnel stability ratio
NSPT
Standard Penetration Test result (blows/300mm)
pF
face resistance stress
PF
face resistance force
Pmisalignment
resistance due to misalignment of pipe string
PF
face resistance force
pF
face resistance stress
PS
skin friction resistance force
ps
skin friction resistance stress
Ptotal
total resistance to jacking force v
Pu
contact force per unit length
q
deviator stress
qt
corrected cone resistance in the cone penetration test
Rf
friction ratio in the cone penetration test
Rinter
interface strength reduction factor in Plaxis software
Tγ
tunnel stability number
u
Pore fluid pressure
w
contact stress due to mass of product pipe
W
contact force per metre due to self-‐weight of product pipe
z0
depth to pipe jack centreline/axis
zw
depth from ground water to pipe jack centreline/axis
α
adhesion factor
β
coefficient characterizing the soil-‐pipe interface
γ
bulk unit weight of soil
γd
dry unit weight of soil
γsat
saturated
γ’
submerged unit weight of soil
γw
unit weight of water or of bentonite slurry
δ, δ’
angle of interface friction (total, effective stress)
Δ
to signify an increase (or decrease) in a quantity
ΔV
magnitude of elastic unloading in the vertical plane
ΔH
magnitude of elastic unloading in the horizontal plane
ε
strain
εa
axial strain
μ
coefficient of friction
ν
Poisson’s ratio
νs
Poisson’s ratio of soil
νp
Poisson’s ratio of product pipes
σT
tunnel internal support pressure
σH
horizontal stress
σm
membrane correction
σV
vertical stress
τ
shear stress
ϕ, ϕ’
angle of internal friction of soil (total, effective stress)
ϕ’cv
constant volume angle of effective friction resistance
ϕ’cv,DS
ϕ’cv determined in direct shear vi
ϕ’cv,TX
ϕ’cv determined in triaxial compression
ψ
angle of dilation
List of abbreviations AVN
automatischer Vortrieb nass (de) / remote controlled, slurry circuit microtunnelling machine
CPT
cone penetration test
DBC
Dublin Boulder Clay
DN
Diamètre Nominal (fr)/nominal diameter
GRP
glass-‐reinforced plastic
MH
manhole
PLT
Point Load Test
PR
Penetration rate in rock (m/hour)
RMR
Rock Mass Rating
RQD
Rock Quality Designation
SPT
Standard Penetration Test
TBM
tunnel boring machine
TCD
Trinity College Dublin
UCS
Uniaxial Compressive Strength
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Publications The following papers have resulted from this research.
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Reilly, C.C. and Orr, T.L.L. (2012), Analysis of interface friction effects on microtunnel jacking forces in coarse-‐grained soils. In: Proceedings of the Bridge and Concrete Research in Ireland (BCRI) Conference. Dublin, 121-‐126.
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Reilly, C.C., McCabe, B.A., Orr, T.L.L. (2012), Analysis of microtunnel jacking forces in alluvium and glacial till in Mullingar, Ireland. In: Proceedings of the ITA/AITES World Tunnelling Congress 2012. Bangkok, Thailand, CD-‐ROM.
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McCabe, B.A., Orr, T.L.L., Reilly, C.C. and Curran, B.G. (2012), Settlement trough parameters for tunnels in Irish glacial tills. In: Tunnelling and Underground Space Technology. 27(1), 1-‐12.
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Reilly, C.C. and Orr, T.L.L. (2011), Microtunnelling -‐ recent experience in Ireland. Proc. International No-‐Dig 2011 Conference & Exhibition, Berlin, 2A-‐4-‐1 – 2A-‐4-‐11.
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Reilly, Ciaran C. (2011), Jacking force measurements at two microtunnelling projects in soft ground in Ireland. Pipeline Industries Guild (Irish Branch) prize-‐winning paper, unpublished.
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Table of contents
Declaration .................................................................................................................................. i Abstract ..................................................................................................................................... iii Acknowledgements ................................................................................................................ iv List of symbols ........................................................................................................................... v List of abbreviations ............................................................................................................ vii Publications ........................................................................................................................... viii 1 Introduction ..................................................................................................................... 14 1.1 Background ............................................................................................................................ 14 1.2 Motivation for the research & research objectives .................................................. 17 1.3 Scope of dissertation ........................................................................................................... 17
2 Literature review ........................................................................................................... 19 2.1 Introduction ........................................................................................................................... 19 2.2 Jacking force in pipe jacking ............................................................................................ 19 2.2.1 Jacking force ...................................................................................................................................... 19 2.2.2 Face Resistance ................................................................................................................................ 21 2.2.3 Frictional Resistance ..................................................................................................................... 21 2.3 Friction .................................................................................................................................... 22 2.3.1 Friction in geotechnical engineering ...................................................................................... 22 2.3.2 Friction theory ................................................................................................................................. 22 2.3.3 Friction in soils and granular materials ................................................................................ 23 2.3.4 Friction in soil-‐structure interfaces ........................................................................................ 28 2.4 Undrained shear strength ................................................................................................. 31 2.5 Use of lubricants in pipe jacking ..................................................................................... 32 2.5.1 Definition of lubrication ............................................................................................................... 32 2.5.2 Lubrication theory .......................................................................................................................... 33 2.5.3 Slurries in geotechnical engineering ...................................................................................... 33 2.5.4 Pipe jacking lubricants ................................................................................................................. 36
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2.5.5 Guidelines for lubricating pipe jacks and usage in the field ......................................... 39 2.6 Modelling the skin friction component of jacking force ......................................... 42 2.6.1 Approaches to modelling the skin friction stress in pipe jacking .............................. 42 2.6.2 Normal stress on underground pipelines ............................................................................ 42 2.6.3 Skin friction resistance models ................................................................................................. 47 2.6.4 Discussion on approaches to modelling skin friction stress ........................................ 57 2.7 Field observations ............................................................................................................... 57 2.7.1 Introduction to field studies carried out .............................................................................. 57 2.7.2 Microtunnelling field research projects ................................................................................ 57 2.7.3 Field case study in a fine-‐grained soil .................................................................................... 60 2.7.4 Field research in Ireland .............................................................................................................. 60 2.8 Physical modelling ............................................................................................................... 60 2.8.1 Advantages of physical modelling ........................................................................................... 60 2.8.2 The physical modelling of interface friction phenomena .............................................. 61 2.8.3 The physical modelling of pipe jacking operations .......................................................... 62 2.9 Summary ................................................................................................................................. 65
3 Field case histories ........................................................................................................ 67 3.1 Introduction ........................................................................................................................... 67 3.2 Pipe jacking and microtunnelling in Ireland .............................................................. 67 3.2.1 Experience of utility tunnelling, pipe jacking and microtunnelling in Ireland .... 67 3.2.2 Segmental utility tunnelling ....................................................................................................... 68 3.2.3 Pipe jacking and microtunnelling ............................................................................................ 68 3.3 Selection of sites ................................................................................................................... 69 3.4 Jacking force measurement and analysis .................................................................... 70 3.4.1 Jacking force measurement ........................................................................................................ 70 3.4.2 Methods of analysis of jacking force ....................................................................................... 71 3.5 Field case histories .............................................................................................................. 73 3.5.1 Presentation of field case histories ......................................................................................... 73 3.5.2 Howth ................................................................................................................................................... 73 3.5.3 Downpatrick ..................................................................................................................................... 77 3.5.4 Mullingar ............................................................................................................................................ 82 3.6 Summary of case history jacking force parameters ................................................. 86 3.6.1 Presentation of summary ............................................................................................................ 86 3.6.2 Effects of lubrication ..................................................................................................................... 86 3.6.3 Operator influence ......................................................................................................................... 86 3.6.4 Pipe surface roughness ................................................................................................................ 87 3.6.5 Curved drives and misalignments ........................................................................................... 87
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3.6.6 Influence of soil type ..................................................................................................................... 89 3.6.7 Relationship between face resistance and skin friction resistance stress ............. 89 3.7 Summary & further remarks ............................................................................................ 93
4 Interface friction reduction with unpressurised lubricants ........................... 95 4.1 Introduction ........................................................................................................................... 95 4.2 Experimental considerations ........................................................................................... 95 4.2.1 Type of soil and variation along the pipeline ..................................................................... 96 4.2.2 Normal stresses acting on the pipeline ................................................................................. 96 4.2.3 Roughness of the pipeline surface ........................................................................................... 99 4.2.4 Rate of advance ................................................................................................................................ 99 4.2.5 Contact conditions between pipe and soil ........................................................................ 101 4.2.6 Overcutting during excavation .............................................................................................. 101 4.2.7 Level of the water table ............................................................................................................. 101 4.2.8 Type and consistency of lubricant ........................................................................................ 102 4.2.9 Duration of stoppages during driving ................................................................................. 103 4.2.10 Pipeline misalignment and joint effects .......................................................................... 103 4.2.11 Boundary effects in the direct shear apparatus ........................................................... 104 4.3 Equipment ............................................................................................................................ 104 4.4 Test materials ...................................................................................................................... 105 4.4.1 Geomaterials .................................................................................................................................. 105 4.4.2 Interface materials ...................................................................................................................... 111 4.4.3 Slurries ............................................................................................................................................. 115 4.5 Testing programme and procedures ........................................................................... 116 4.5.1 Testing programme .................................................................................................................... 116 4.5.2 Testing procedures ..................................................................................................................... 116 4.6 Results .................................................................................................................................... 118 4.6.1 IGB sand ........................................................................................................................................... 119 4.6.2 Banagher sand ............................................................................................................................... 127 4.6.3 Kaolin clay ....................................................................................................................................... 134 4.6.4 Cloughfin clay ................................................................................................................................ 137 4.6.5 Summary of results ..................................................................................................................... 138 4.7 Summary & further remarks .......................................................................................... 140
5 Interface friction reduction with pressurised lubricants – experimental apparatus and procedures ............................................................................................... 142 5.1 Introduction ......................................................................................................................... 142 5.2 Experimental considerations ......................................................................................... 142
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5.2.1 Requirements for the design of a physical model .......................................................... 142 5.2.2 Replicating field conditions in a physical model ............................................................ 143 5.2.3 Implementing a physical model in the laboratory ........................................................ 143 5.3 Testing apparatus and materials .................................................................................. 145 5.3.1 Development of the test apparatus ...................................................................................... 145 5.3.2 Triaxial testing in geotechnical engineering .................................................................... 145 5.3.3 Shear stress in triaxial testing ................................................................................................ 146 5.3.4 Geomaterials .................................................................................................................................. 147 5.3.5 Development of preformed failure plane triaxial testing apparatus ..................... 152 5.3.6 Lubricant slurries ........................................................................................................................ 158 5.4 Test procedures .................................................................................................................. 159 5.4.1 General procedures ..................................................................................................................... 159 5.4.2 Unlubricated tests ....................................................................................................................... 161 5.4.3 Lubricated tests ............................................................................................................................ 161 5.5 Corrections ........................................................................................................................... 164 5.5.1 The need for corrections to the measured axial force ................................................. 164 5.5.2 Area corrections ........................................................................................................................... 165 5.5.3 Membrane corrections .............................................................................................................. 167 5.5.4 Corrections applied to test results ....................................................................................... 170 5.6 Summary ............................................................................................................................... 170
6 Interface friction reduction with pressurised lubricants – results ............. 171 6.1 Introduction ......................................................................................................................... 171 6.2 Behaviour of the novel triaxial interface test apparatus ...................................... 171 6.2.1 Interpretation of the model behaviour & interface strength .................................... 171 6.2.2 Interface friction with IGB sand ............................................................................................ 174 6.2.3 Interface friction with Banagher sand ................................................................................ 175 6.2.4 Numerical modelling of preformed failure plane behaviour .................................... 176 6.3 Effects of lubricant injection .......................................................................................... 181 6.3.1 Static injection ............................................................................................................................... 182 6.3.2 Dynamic injection at a specified point in time ................................................................ 192 6.3.3 Constant pressure lubricant injection ................................................................................ 193 6.4 Discussion ............................................................................................................................. 194 6.4.1 Commentary on lubricated test results .............................................................................. 194 6.4.2 Comparison of results with other physical modelling programmes ..................... 199 6.4.3 Comparison of results with the findings from field studies ...................................... 200 6.4.4 Recommendations towards practice ................................................................................... 201 6.5 Summary ............................................................................................................................... 201
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7 Pipe jacking and microtunnelling in hard rock ................................................. 203 7.1 Introduction ......................................................................................................................... 203 7.2 Jacking forces and rates of advance during microtunnelling in hard rock .... 204 7.2.1 Hard rock microtunnelling compared to soft soil microtunnelling ....................... 204 7.2.2 Rock materials and classifications as applied to TBM tunnelling ........................... 204 7.2.3 Face resistance .............................................................................................................................. 206 7.2.4 Friction resistance ....................................................................................................................... 207 7.2.5 Advance rate and penetration rate ...................................................................................... 208 7.3 Field case histories in hard rock ................................................................................... 209 7.3.1 Selection of sites ........................................................................................................................... 209 7.3.2 Tullamore ........................................................................................................................................ 210 7.3.3 Gilford ............................................................................................................................................... 218 7.4 Discussion ............................................................................................................................. 223 7.4.1 General comments on results ................................................................................................. 223 7.4.2 Jacking force predictions in hard rock ................................................................................ 224 7.4.3 Comparisons to published results ........................................................................................ 225 7.5 Summary ............................................................................................................................... 227
8 Conclusion ...................................................................................................................... 228 8.1 Introduction ......................................................................................................................... 228 8.2 Conclusions .......................................................................................................................... 229 8.2.1 Field case histories ...................................................................................................................... 229 8.2.2 Physical modelling ...................................................................................................................... 230 8.3 Summary ............................................................................................................................... 233 8.4 Recommendations for future research ....................................................................... 234
References ............................................................................................................................. 235 Appendix A – Micro concrete mix design .................................................................... 244 Appendix B – Hydraul-‐EZ technical data sheet ......................................................... 245 Appendix C – MX Polymer technical datasheet ......................................................... 246 Appendix D – Triaxial testing procedure .................................................................... 247 Appendix E – Table of triaxial interface shear testing results ............................. 250
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Chapter 2
Literature review
1 Introduction
1.1 Background As we are living in a rapidly urbanising world, with more people now living in metropolitan areas than rural, trenchless technologies such as pipe jacking and microtunnelling are becoming more and more important to towns and cities seeking to retain and improve the standards of utilities available to their citizens in a sustainable manner. Pipe jacking or microtunnelling are construction techniques that allow for the underground installation of pipelines from a starting shaft to a reception shaft some distance away, with a minimum of vertical intrusions to the surface. These methods allow for the installation of pipelines in diameters ranging from 300mm to 3000mm. Typical uses include water, gas or oil supply pipelines, waste water collection systems and electrical and telecoms cable ducting. Microtunnelling is a special case of the pipe jacking technique, although the terms are used almost interchangeably. In microtunnelling, excavation at the front of the pipeline is carried out using a remote-‐controlled mechanised shield with a cutting wheel at the front and spoil removal to the surface via a slurry circuit or a conveyor system. Microtunnelling is currently a very advanced art with high levels of automation and monitoring available for many of the processes involved, and, for most of the time, personnel are not located in the pipeline. A typical microtunnelling arrangement is shown in Figure 1.1. Traditional pipe jacking, on the other hand, requires the presence of personnel within the pipeline and the shields used may consist of open-‐faced enclosures designed for hand digging or fitted with mechanical excavation systems operated by personnel protected within the shield. The choice of technique will often come down to the type of ground that must be excavated, the general classifications for the purposes of equipment selection being soft soil, hard soil/soft rock and hard rock. Remote-‐controlled microtunnelling tends to be preferred for longer lengths of pipeline and the more challenging, i.e. very soft or very hard, types of soil. A conceptual example of a project utilising a hard rock microtunnelling machine to install pipes on a curved alignment under the seabed is shown in Figure 1.2.
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Chapter 2
Literature review
Figure 1.1 -‐ Schematic of slurry microtunnelling operation (Source: Iseki Microtunnelling)
Figure 1.2 -‐ Installation of subsea pipeline or outfall using curved microtunnelling technique (Source: Herrenknecht AG)
A large force, the jacking force, is required to advance the microtunnelling machine or pipe jacking shield and the string of product pipes behind it from the starting shaft through the ground to the reception shaft. The force required must be developed by the jacking
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Chapter 2
Literature review
apparatus and resisted by a thrust structure in the starting shaft, usually a thrust wall that transfers the force to the ground. The product pipes chosen for the pipe jacking operation, whether concrete, steel, ceramic or glass-‐reinforced plastic, must be capable of sustaining the maximum expected force without damage to the pipe. Therefore it is essential that the likely jacking forces for given conditions can be estimated in advance of a project. The total jacking force must overcome the resistances at the face of the shield and also skin frictional resistance developed over the surface of the shield and the pipeline, as shown in Figure 1.3.
Figure 1.3 -‐ Face resistance, skin frictional resistance and total jacking forces
The microtunnelling machine or pipe jacking shield typically excavates an opening with a slightly larger diameter than the diameter of the product pipe to be installed. This “overcut” or “overbreak” helps to reduce the skin frictional resistance acting on the pipeline and facilitates the steering of the microtunnelling machine. Bentonite-‐ or polymer-‐based lubricant slurries may be injected into this overcut through ports set in the pipes, as shown in Figure 1.4, to further reduce the skin frictional resistance. This thesis is concerned with the influence of these lubricant slurries in reducing the skin frictional resistance, and hence the total jacking force, in pipe jacking.
Figure 1.4 – Typical arrangement of lubrication ports (Pipe Jacking Association, 2007)
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Chapter 2
Literature review
1.2 Motivation for the research & research objectives The mechanisms of skin friction reduction due to the injection of lubricant slurries into the overbreak during pipe jacking are not well documented or understood. Three main mechanisms have been suggested for the reduction in skin frictional resistance due to the injection of bentonite-‐ or polymer-‐based slurries, as follows: 1. Conventional boundary lubrication reducing the sliding resistance of the interface between the soil and the pipe. 2. Rendering of the pipeline fully or partially buoyant in the viscous slurry. 3. Reduction of the radial effective stress acting on the pipeline. The present research seeks to clarify the extent of the contribution of each of these mechanisms to the measured jacking force reductions observed in the field through the collation and analysis of field case history data and the physical modelling of the phenomena in the laboratory. It was sought to investigate in particular whether the lubricant slurries actually “lubricated” the interface, or if some other mechanism, such as radial effective stress reduction, was more significant. This was considered important as the conventional approaches to pipe jacking lubrication in the field focus on the control of the volume of lubricant injected, while it was considered that the control of lubricant injection pressure might have more merit. This research is topical, since due to increasing environmental concern, ever more complex microtunnelling projects are conceived in fields such as long sea outfalls, seawater abstraction intakes and ambitious undersea oil and gas pipelines. These demands are broadening the remit of microtunnelling from the traditional urban setting towards deeper and longer drives. Such extremely long distance pipe jacking is still very challenging, and a thorough understanding of the mechanism of skin friction reduction due to the injection of lubricant slurries is vital to solving the problems that will be faced.
1.3 Scope of dissertation The thesis layout is as follows: Chapter 2:
A general introduction to jacking force in pipe jacking and microtunnelling is given, along with an introduction to the use of lubricants to reduce skin frictional resistances. Previous research in the area is introduced and reviewed, focussing on empirical, analytical and physical models used to analyse and predict the skin frictional resistance component of jacking force
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Chapter 2
Literature review
with and without lubrication. Field studies carried out by others are appraised, with an emphasis on studies outlining the impact of lubrication on jacking forces. Chapter 3:
Field case histories for three microtunnelling projects in soft soil collated and analysed by the author are presented and discussed. The jacking forces are analysed for each drive and the impact of soil properties, operational factors and lubrication on the measured jacking forces is reviewed. The results obtained through back analysis of the field case history data are compared to results found in the literature.
Chapter 4:
The use of the direct shear soil testing apparatus to investigate the skin frictional resistance between a number of different soils and a rough concrete interface, which was similar in roughness to a concrete jacking pipe, is described. The influence of unpressurised pipe jacking lubricants applied to the interface was examined, and the lubrication mechanisms that occurred were identified.
Chapter 5:
The development of a novel triaxial testing procedure that allowed for the carrying out of interface friction testing is described. Furthermore, it is explained how the apparatus allowed for the injection of pipe jacking lubricants under pressure into the interface between a soil and a rough concrete test specimen.
Chapter 6:
The results obtained from experimental testing using the novel triaxial testing apparatus incorporating a soil specimen and a concrete specimen are presented and analysed. Testing was carried out to simulate the injection of lubricants under pressure during pipe jacking and the resulting effects and mechanisms of lubrication are identified and examined.
Chapter 7:
Recent case histories of two microtunnelled pipelines in hard rock are presented, along with an analysis of the jacking forces, advance rates and operational factors encountered in each. The effectiveness of lubrication in hard rock pipe jacking situations is discussed.
Chapter 8:
Finally, the main findings and conclusions of the field and experimental work are summarised, highlighting the practical applications of the research and suggesting areas where further work may be recommended.
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Chapter 2
Literature review
2 Literature review
2.1 Introduction The research described in this thesis concerns the skin friction resistance generated during the installation of pipe-‐jacked underground pipelines and how this skin friction resistance may be reduced through the injection of lubricant slurries into the ground around the pipeline during pipe jacking. In this chapter, the background to this research is set out. The concept of jacking force is described and a distinction between total jacking force and skin friction resistance is made (§2.2). The theory of friction is discussed and its relevance to soil mechanics is explained (§2.3), while the undrained shear strength and interface shear strength of fine-‐grained soils is also introduced (§2.4). Lubrication mechanisms and the use of slurries as lubricants in pipe jacking are described and the manner in which pipe jacking lubricants are applied in the field is set out (§2.5). The analytical models commonly used for predicting or back analysing the skin friction component of jacking force are examined (§2.6). Some of the field research carried out to date into jacking forces in pipe jacking is reviewed (§2.7), while work done to physically model the skin frictional resistance is discussed (§2.8). Finally, the chapter is summarised (§2.9).
2.2 Jacking force in pipe jacking 2.2.1 Jacking force The jacking force is that force which must be applied to the rear of the pipe string to advance the product pipes through the ground. The prediction of jacking forces in advance of a project is difficult but important. It is difficult due to the complex soil-‐structure interactions present, and important because it must be ensured that the specified materials and equipment are appropriate to safely complete the required drive distance. According to Stein & Beckmann (2011), there are three sources of resistance likely to act on the pipe jacking shield or microtunnelling machine and jacking pipes as they are advanced through the ground: 1. Face resistance (force: PF and stress: pF), which is the resistance that the face of the pipe jacking shield or microtunnelling machine must overcome as the soil at the face is excavated. 2. Skin frictional resistance (force: PS and stress: pS), which is due to the skin friction developed over the external surface area of the pipe jacking shield or microtunnelling machine and jacking pipes as they are pushed into the ground. It is
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Chapter 2
Literature review
useful to note at this stage that this resistance is almost always frictional in nature, rather than adhesive (Norris, 1992). 3. Any extra resistance caused by unintentional misalignments and corrections there of, or intentional curvature introduced during the pipe jacking process (force: PMisaliggnment and stress: pMisalignment). Once the soil type and the operational parameters remain the same, the face resistance should remain uniform throughout the drive, but as the length of pipeline increases, so too will the skin friction resistance and misalignment resistance. This is shown in Figure 2.1. While it is relatively easy to mathematically model the face resistance and skin friction resistance, misalignment resistance is much harder to predict and model (Sugimoto and Asanprakit, 2010), varying greatly depending on the specific parameters of a particular drive. For these reasons, misalignment resistance will not be studied in detail here. As most of the case histories presented throughout this research are of straight drives and it is universally the case that attempts are made by the contractors to maintain trueness in these drives, it is considered appropriate to allow for misalignment resistance within the analysis of face and skin frictional resistance.
Figure 2.1 -‐ Illustrative jacking force plot against length driven, showing hypothetical contributions from each type of resistance
If it is assumed that misalignment resistance may be accounted for within the face and the skin frictional resistance, the total resistance to be overcome by the jacking force can be expressed as the sum of the face resistance force, PF and skin friction resistance force, PS: Ptotal = PF + PS
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Equation 2-‐1
Chapter 2
Literature review
2.2.2 Face Resistance The resistance to advance of the face of a pipe jacking shield or microtunnelling machine through the ground is a function of the ground strength in front of the face and the speed at which the microtunnelling machine is advanced. While with open face pipe jacking the greatest risk will be the collapse of the face inwards, with microtunnelling machines there is also the possibility of exerting excess force on the ground in front of the machine. As microtunnelling machines are designed and operated so as to achieve a pressure balance at the face, the face resistance will vary between the active and passive earth pressure acting on the face of the machine. Ideally the microtunnelling machine will be driven as fast as possible without stalling due to lack of torque. If rate of advance is too slow, there is increasing the risk of material sloughing into the excavation and subsequent surface settlement, while too fast a rate may cause heave (Staheli, 2006). The solids and liquids (groundwater and slurry) present must be also controlled so as to avoid both soil uplift and surface settlement, as shown in Figure 2.2 and Figure 2.3.
Figure 2.2 – Forces acting on the face of a
Figure 2.3 – Possible deformations caused by
microtunnelling machine (Stein, 1989)
face pressures in pipe jacking (Stein, 1989)
2.2.3 Frictional Resistance As shown in Figure 2.1 the skin frictional resistance force generally makes up the larger part of the total resistance to jacking force. As the length of drive gets longer, so does the overall frictional resistance, and it is ultimately the frictional resistance that determines the practical lengths of individual drives. Several factors affect the skin frictional resistances during pipe jacking including (Marshall, 1998; Chapman and Ichioka, 1999):
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Type of soil and variation along the pipeline
•
Normal stresses acting on pipeline 21
Chapter 2
Literature review
•
Roughness of pipeline surface
•
Rate of strain
•
Contact conditions between pipe and soil
•
Overcut during excavation
•
Position of water table
•
Lubrication
•
Duration of stoppages, for example breakdowns, overnight or weekends
•
Pipeline misalignment and/or intentional curvature.
•
Changes to soil characteristics due to dynamic loading and vibrations.
This list of factors is widely accepted in the literature to hold true for the modelling of frictional resistance and will be used later in this thesis to provide a framework for investigations into skin friction resistance.
2.3 Friction 2.3.1 Friction in geotechnical engineering Since this research is primarily concerned with the effects of frictional resistances to movement, it is useful to review the theory of friction between solid bodies, and to discuss frictional contact in soils. By way of comparison, cohesion or adhesion in soils will be briefly commented upon. Since a great many textbooks expand on these topics in detail, this section will be brief.
2.3.2 Friction theory “Friction” is intuitively the force that exists between two objects in contact that resists relative movement between them. Leonardo da Vinci (1452-‐1519) experimentally verified the two basic laws of friction, that the frictional resistance is proportional to the load applied and independent of the area of contact of the sliding surfaces. In 1699, Amontons found similarly, and incorrectly proposed that friction was due to interlocking asperities on the surfaces in contact and the frictional force was always equal to one-‐third of the normal load. Coulomb in 1781 made a distinction between static and dynamic friction and showed that dynamic friction was lower than static friction. Coulomb was unsure whether friction was due to adhesion between molecules on each surface or the interlocking of surface asperities (Bowden and Tabor, 1958), however the Coulomb model of friction has persisted: 𝐹 = 𝜇𝑁
22
Equation 2-‐2
Chapter 2
Literature review
where F is the force resisting sliding on a plane, N is the normal force acting perpendicularly to the plane and μ is the coefficient of friction, or the ratio of the resisting force to the normal force acting n the plane. It is convenient to let μ = tan ϕ, where ϕ is termed the angle of friction, and to express the forces N and T as the stresses σ and τ acting over the area of contact: 𝜏 = 𝜎 𝑡𝑎𝑛𝜙
Equation 2-‐3
A common representation of friction, the frictional slip of a block on a plane, is shown in Figure 2.4.
Figure 2.4 -‐ Frictional slip of a block on a plane
Much of our modern understanding of frictional processes comes from Bowden and Tabor’s classic work in friction theory (1958) which deals in the main with metals. Frictional stresses are generated by chemical bonding or “welding” in and around areas of real contact, i.e. areas where tiny asperities on the surface of materials come into contact, areas that are many times smaller than the geometric area of contact. This area of true contact is found to be roughly proportional to the applied load, which fits in with the previously mentioned observations made by Da Vinci and Coulomb many centuries ago.
2.3.3 Friction in soils and granular materials 2.3.3.1 Friction mechanisms Just as adhesive bonding prevents relative movement between two flat surfaces, the movement of two soil grains against one another is similarly resisted at points of real contact, as is the movement of assemblies of soil grains, as illustrated in Figure 2.5. Equation 2-‐3 equally holds for the resistance to movement in all of these situations.
23
Chapter 2
Literature review
Figure 2.5 -‐ Friction between sliding bodies (Bolton, 1979)
While metals can flow plastically and “weld” together at the areas of real contact, crystalline solids like many soils cannot and hence form junctions which are weaker. Minerals and rocks that normally form granular soil particles have distinct inter-‐particle sliding friction angles, depending on factors such as surface micro roughness, durability, texture and hardness of the surface and by the crystal structure and bonding characteristics of the mineral crystals. Values of inter-‐particle sliding friction range from 20° to 40°, with more common materials having values in the range 25° to 35° (Terzaghi et al., 1996).
2.3.3.2 Effective stress The presence of water within a soil does not affect the coefficient of friction between the grains of a soil, but the pressures that exist in the water will have profound effect on the soil behaviour. This behaviour was described by the Principal of Effective Stress (Terzaghi, 1943), where the total normal stress on a plane, σ, is equal to the sum of the pore water pressure, u, and the effective stress acting through the soil skeleton only (σ’). Thus, Equation 2-‐3 becomes: 𝜏 = 𝜎 − 𝑢 𝑡𝑎𝑛𝜙′ = 𝜎′ 𝑡𝑎𝑛𝜙′
Equation 2-‐4
where τ’ and ϕ’ are the effective shear stress and effective angle of shearing resistance respectively.
2.3.3.3 Shear strength and failure in soils The strength of a soil is not a unique value but rather depends very much on the conditions under which the soil is loaded. A number of failure criteria exist, that allow for the prediction of the strength of a soil under certain specified conditions once strength tests have been carried out on a representative sample of the soil under known conditions. For drained conditions, where excess pore water pressure is not developed, the Coulomb failure criterion applies for non-‐cemented soils. It is similar in formulation to Equation 2-‐4, but
24
Chapter 2
Literature review
using the parameters that correspond to a state of shear failure, i.e. slippage along the plane S-‐S in Figure 2.5, τf’ and σf’: 𝜏! = 𝜎! ′ 𝑡𝑎𝑛𝜙′
Equation 2-‐5
This condition may be visualised using the results of a series of triaxial tests on the same soil, where the major and minor principal stresses at failure are used to plot Mohr circles of stress and a line through the origin and produced tangential to these circles represents the soil strength envelope, as illustrated in Figure 2.6. The soil cannot attain a stress condition above this envelope. The compressive sense of the Mohr circles is of interest only, and so the tensile region is removed in the presentation of Mohr circles of stress for soils.
Figure 2.6 -‐ Coulomb shear strength envelope
The Coulomb failure criterion is suitable for coarse-‐grained soils where the behaviour is purely frictional. In some soils, particularly fine-‐grained soils, a drained “cohesion” intercept is present, where the shear stress is positive when the normal stress is zero. Various explanations exist for this phenomenon, including real cohesion and cementing, but these will not be dealt with here. The Mohr-‐Coulomb criterion may be used to define the failure conditions of such a soil and Equation 2-‐5 may be re-‐written to take account of a drained cohesion intercept c’: 𝜏! = 𝑐 ! + 𝜎! ′ 𝑡𝑎𝑛𝜙′
Equation 2-‐6
Using hypothetical test results again, the Mohr-‐Coulomb failure criterion may be illustrated as shown in Figure 2.7.
25
Chapter 2
Literature review
Figure 2.7 -‐ Mohr-‐Coulomb shear strength envelope
The shear strength of frictional soils under drained loading conditions may be partially described by these two models. In order to more fully model the shearing behaviour of certain soils, other concepts must be taken into account, for example dilation and peak strength. These will be briefly introduced in the following section.
2.3.3.4 Peak strength and dilatancy Critical state theory (Muir Wood, 1990) inextricably links shearing and volume change in soils. When a loose sand shears it tends to contract, whereas a dense sand will dilate, or expand in volume until it reaches a critical state of constant volume shearing. This phenomenon may be demonstrated quite simply in the shear box. When a dense sand dilates, the lid of the shear box moves upwards at the angle of dilation ψ . There are a number of explanations for this behaviour. This occurs as the interlocking particles of sand rearrange themselves as they move past each other. The conventional approach commonly seen in text books is that of the saw tooth (Bolton, 1979), where the particles must rise up microscopic shear planes inclined at an angle of ψ to the horizontal in order to make progress. This saw tooth analogy, as shown in Figure 2.8 below, can only be relevant for very small strains, as when one tooth crosses the opposite tooth and interlocks again, the net volume change is zero, or for very fast strain rates where the saw tooth surfaces slide so quickly that only the tips touch.
Figure 2.8 -‐ Saw tooth analogy for dilation (Powrie, 2002)
26
Chapter 2
Literature review
More recently it has been suggested that dilatancy and critical state in granular materials could be explained by an evolving frequency of microscopic topological events (flips) and elastic relaxation of the granular particles (Mesarovic et al., 2012), each giving rise to net volume increases until a strain of approximately 5% is reached. Figure 2.9 shows a particular “cell” undergoing a volume increase (dilation), followed by contraction.
Figure 2.9 – Generic flipping mechanism (Mesarovic et al., 2012).
From left to right, as the cluster is vertically compressed, the status of neighbouring particles changes. Solid lines indicate particles in contact, whereas dashed lines indicate nearest neighbours not touching. The volume occupied by the cluster first increases by Δε+ then decreases by Δε-‐. It was found that at each time increment, only a small fraction of cells in a sample are undergoing flipping, and in the next increment a completely different set of cells are topologically translating, allowing for the net volume increase observed during the shearing of dense sands. This property of dilatancy allows dense sands undergoing shearing to exhibit a higher peak angle of shearing resistance before reaching a lower, critical state, angle of shearing resistance. The peak angle of resistance, ϕ’peak, is a transient quantity whereas the critical state angle of shearing resistance ϕ’ cv is a fundamental parameter of a material (Powrie, 2002). For sands, this critical state angle of friction is generally termed the constant volume angle of shearing resistance, while in fine-‐grained materials, another strength, the residual strength, may often be quoted as the minimum strength of the material after relative movement has caused clay particles to become aligned along a preferred failure plane (Lupini et al., 1981). The concept of critical state shearing is illustrated in Figure 2.10, where the typical shearing response of three different classifications of soil are shown (Budhu, 2011). Type I soils are loose sands, normally consolidated and lightly overconsolidated clays, Type II soils are dense sands and heavily overconsolidated clays and Type II-‐A soils are clays in which the particles may become aligned along a preferred failure plane, which allows for a lower residual shear strength.
27
CHAPTER 10 SHEAR Chapter 2 STRENGTH OF SOILS
Literature review
τ
Shear stress
τp
Peak Type II—dense sands and overconsolidated clays
Critical state
τcs
Type II–A soils
τr
Type I—loose sands, normally consolidated and lightly overconsolidated clays Shear strain, γzx (a) εz
Compression
e
Void ratio
Type I soils
Vertical strain
Type I soils
ecs Shear strain, γzx
Expansion
264
Critical void ratio
Type II soils Shear strain, γzx
∆εz
(c)
Type II soils
∆γzx – ∆εz –d εz tan α = ––––– = ––––– ∆γzx d γzx (b)
FIGURE 10.3
Response of soils to shearing.
Figure 2.10 -‐ Shearing behaviour of soils in terms of (a) shear stress (b) vertical strain and • Compress initially (attributed to particle and2011) then expand, that is, they become looser (c) void adjustment) ratio (Budhu, (Figure 10.2c and Figure 10.3b, c) until a critical void ratio (the same void ratio as in Type I soils) is attained.
The critical state shear stress is reached for all soils when no further volume change occurs under 2.3.4 Friction n soil-‐structure interfaces continued shearing. iWe will use the term critical state to define the stress state reached by a soil when
2.3.4.1 Interface friction shearing mechanism There has been much research on the shearing mechanism between soil and construction materials (Potyondy, 1961; Uesugi and Kishida, 1986b; Uesugi and Kishida, 1986a; Uesugi et al., 1988; Uesugi et al., 1989; Uesugi et al., 1990; Gómez et al., 1999; O'Kelly and Naughton, 2008). While shearing behaviour may be modelled using the Coulomb or Mohr Coulomb FIGURE 10.4 Radiographs of shear bands a densethe fine angle sand (the failure criteria introduced in section 2.3.3.3, in in general of white internal frictional
circles are lead shot used to trace internal displacements; white lines are
resistance ϕ ’ is sbands). ubstituted for by an 1979.) angle of interface friction resistance δ’, which will vary shear (After Budhu, up to a maximum value of ϕ’ depending on interface contact conditions. An illustration of frictional contact between a soil and an interface, in this case a concrete jacking pipe, is shown in Figure 2.11.
28
Chapter 2
Literature review
Figure 2.11 -‐ Frictional forces (T,N) and stresses (τ,σ) acting on a jacking pipe in contact with in-‐situ soil
The main factors shown to effect interface friction behaviour under direct loading are interface roughness, soil density, particle angularity and normal stress across the interface. A variety of test methods are used in these works and elsewhere to assess the interface behaviour, including direct shear (Potyondy, 1961; Zong-‐Ze et al., 1995; Iscimen, 2004; Liu et al., 2005; O'Kelly and Naughton, 2008), simple shear (Uesugi and Kishida, 1986b; Uesugi and Kishida, 1986a; Kishida and Uesugi, 1987; Uesugi et al., 1988; Uesugi et al., 1989; Uesugi et al., 1990), ring shear (Tan et al., 1998; Hungr and Morgenstern, 1984) and annular shear apparatus (Kishida and Uesugi, 1987).
2.3.4.2 Roughness While a surface may look flat and smooth and feel smooth to the touch, under a microscope all surfaces have a complex structure with irregularities introduced by material structure, coatings applied or the manufacturing processes. Surface roughness or texture often affects the functionality of a surface. In geotechnical engineering, the surface roughness of an interface will have a significant effect on the resistance to shearing between a soil and an interface (Potyondy, 1961; Uesugi and Kishida, 1986b). There are two classes of parameters commonly used for the measurement of surface roughness, profile and area. Profile parameters are parameters based on measurements made over a line of a certain length, the gauge length, while area parameters are derived from measurements made over a certain area (British Standards Institute, 2010). Line measurements are typically made using a moving mechanical stylus that physically makes contact with the surface, while area measurements may be made using more modern optical instruments, which do not make contact with the surface. In common, these measurements are used to produce a surface topography resulting from the intersection of the real surface with a specified plane, usually the plane that lies nominally parallel to the real surface. Linear profiles may be extracted from area profiles if necessary.
29
Chapter 2
Literature review
In geotechnical engineering practice, line profile surface roughness measurements are common (Uesugi and Kishida, 1986b; Iscimen, 2004; Knapett and Craig, 2012). The most commonly quoted profile surface roughness parameters in the literature are Ra and Rmax (also written Rz). Ra is the arithmetic mean deviation of the absolute of the ordinate values over an evaluation length L: 𝑅! =
1 𝐿
𝑧(𝑥) 𝑑𝑥
Equation 2-‐7
!
Rmax is the maximum height difference recorded over the profile, i.e. the distance between the maximum peak height and the minimum valley depth from the plane of measurement over the evaluation length. Uesegi et al. (1990) performed simple shear tests on air-‐dry sand and concrete and found that peak angle of interface friction mobilised between the soil and the concrete is highly dependent on the roughness of the concrete and the mean particle size (D50) of the sand. The normalised roughness parameter, Rn, was introduced, which relates the roughness of the interface surface to the particle size distribution of the soils shearing against it as follows:
𝑅! =
!!"# (!≈!!" ) !!"
Equation 2-‐8
where Rmax(L≈D50) is the height difference between the highest peak and the lowest valley taken along a surface profile with a length L equal to the D50 value of the soil under investigation. A relation between the coefficient of friction, μ, which is the tangent to the angle of interface friction δ, is shown in Figure 2.12. It is seen that as the normalised roughness increases, a “critical” roughness is reached after which the coefficient of friction increases no more. This critical roughness is the roughness value at which the limiting interface shear resistance changes from being a function of the interface angle of friction, δ’, to the angle of shearing resistance of the sand itself, ϕ’, i.e. the surface of shear failure moves away from the interface and into the soil body. In practice, the rougher the interface and/or the smaller the average particle size of a soil, the greater the interface shearing resistance becomes until the limiting value is reached.
30
Deep foundations
Chapter 2
of best fit is then tan δ′. FigureLiterature review effective stress; the gradient of the line 9.8 shows data from interface shear tests collated from the literature (Uesugi and Kishida, 1986; Uesugi et al., 1990; Subba Rao et al., 1998; Frost et al., 2002) for a range of common pile materials in which δ′ is expressed in terms of the angle of shearing resistance of the soil. These data are plotted against the parameter Ra/D50, where Ra is the average height of the asperities on the surface of the pile material which gives the material its roughness and D50 is the mean particle size of the soil (see Chapter 1). It should be noted that mild steel oxidises (rusts) in the ground, so δ′/b′ is usually limited to a minimum of 0.5. The concrete values given in Figure 9.8 are for precast concrete (i.e. for driven piles); for concrete cast in-situ the roughness will be much larger, and δ′ = b′ is usually assumed. In fine-grained materials under drained conditions, K ≈ K0 and δ′ = b′ is usually assumed. Alternatively, the parameters K and tan δ′ may be lumped together into a single factor β = Ktan δ′, i.e. (9.10) The determination of shaft friction in drained soil is often referred to as the β-method in pile design. Burland (1993) showed that β correlates linearly with the yield stress ratio with a surprisingly small amount of scatter, as shown in Figure 9.9 (data points represent non-displacement piles). A best-fit line to this data gives (9.11)
Equation 9.11 has been found to provide reasonable estimates of shaft capacity for both displacement and non-displacement piles. In the case of underreamed piles, as a result of settlement there&is Kaishida, possibility that a small gap will Figure 2.12 -‐ "Critical" roughness on the sand-‐concrete interface (Uesugi 1990) develop between the top of the under-ream and the overlying soil. Accordingly, no skin friction should be taken into account below a level 2D0 above the top of the under-ream, and base resistance should be determined if the base not embedded Figure 9.5 case (b) with d = 0, spresent cNc = 6.2). a, Based on a review of a as number of isother studies, (i.e. Knapett and Craig (2012)
perhaps more practical, graph of the interface friction ratio against normalised average
Pile resistance and limit state design
surface roughness (Figure 2.13), showing similar behaviour with increasing surface
The resistance of a pile is the sum of its base and shaft capacities. In the limit state design framework,
the rcombined compressive resistance may be factored using a partial factor γRC to obtain the roughness and/or educing a(total) verage particle size. design resistance, i.e.
1 0.8 0.6 0.4 Concrete (precast) Steel
0.2 0 0.00001
Timber 0.0001
0.001
0.01
0.1
1
Ra/D50
Figurea9.8 friction angles `a for various construction materials. Figure 2.13 -‐ Friction ngle Interface ratio against average roughness normalised by soil mean particle size D50 (Knapett & Craig, 2012)
2.4 Undrained shear strength While field studies have shown that the soil around advancing jacking pipes generally behaves in a drained manner (Norris, 1992), Anagnostou and Kovari (1996) give a range of conditions at which it is predicted soil around a tunnel will behave in the manner of an undrained soil. It is suggested that when the soil affected has a permeability lower than
31
335
Chapter 2
Literature review
roughly 10-‐7 to 10-‐6 m/s and when the penetration rate is faster than roughly 1.67 or 16.7 mm/min, then the soil around the tunnel will behave as an undrained soil. It is useful then to consider the undrained shear strength of a soil, i.e. the resistance to shearing while significant excess pore pressures are allowed to build up within the soil. The undrained shear strength of a soil is denoted cu, which is a total stress parameter, and is described by the Tresca failure criterion: 𝜏! = 𝑐!
Equation 2-‐9
The Mohr circles of stress at failure of three hypothetical samples of an identical fully saturated soil are shown in Figure 2.14, and illustrate that the shear stress at failure will remain uniform despite changes in total normal stress.
Figure 2.14 -‐ Tresca (undrained) shear strength envelope
In undrained soil structure interaction problems, the mobilised shear stress resisting movement along the interface, τint, is related to the undrained shear strength through an adhesion factor α (Knapett and Craig, 2012): 𝜏!"# = 𝛼𝑐!
Equation 2-‐10
The adhesion factor α may range from zero to unity, with a perfectly rough interface having an adhesion factor of unity.
2.5 Use of lubricants in pipe jacking 2.5.1 Definition of lubrication The skin friction resistance between a moving pipe and the surrounding soil depends on the effective normal stress acting on the pipe and the angle of effective interface shearing resistance between the pipe and the soil, as shown in Figure 2.11. If the frictional stress is to be reduced, a reduction in the effective normal stress acting on the pipeline or the angle of
32
Chapter 2
Literature review
interface shearing resistance is required. Lubrication, from the Latin lubricus, “slippery”, is the common term for the action of minimising friction and allowing smooth movement, and the manner in which lubrication may be applied to pipe jacking is discussed in this section.
2.5.2 Lubrication theory In ideal fluid lubrication, surfaces moving relative to each other are entirely separated by a layer of lubricant of appreciable thickness and the resistance to motion is due only to the viscosity of the lubricant layer. In practice, and particularly where speeds of motion are low and loads are high, this situation is unlikely and instead lubrication will be of the film or boundary type (for the purposes of this research, film lubrication and boundary lubrication may be considered as synonymous). Here, lubricant films of molecular thickness separate the surfaces, and the efficiency of lubrication will have little to do with the bulk viscosity of the lubricant and more to do with the properties of the sliding surfaces themselves (Bowden and Tabor, 1958). As for unlubricated friction, the frictional resistance is proportional to normal stress except for extremely low stresses where frictional resistance is found to increase with reducing stress, and the efficiency of film lubrication is generally independent of speed of movement for sensible speeds (Bowden and Tabor, 1958).
2.5.3 Slurries in geotechnical engineering Mineral slurry, chiefly based on bentonite clay, has been used during construction to stabilise drilled shafts in coarse-‐grained soils since the late 1940s (Majano et al., 1994), with polymer slurries popular recently due to environmental constraints affecting the use of mineral slurries. One of the primary requirements of a slurry is the effective formation of a filter cake or mud cake, which is a relatively impermeable region of gelled slurry particles that forms, under the influence of a pressure gradient, on the surface of and to a certain depth within a permeable soil (Figure 2.15). This filter cake of solid particles enables the transfer of positive effective stress to the soil particles and hence helps to maintain the stability of, for example, a borehole.
33
Chapter 2
Literature review
Figure 2.15 -‐ Filter cake formation in soil (Washbourne, 1986)
most typical material used as a slurry base is bentonite, which is the common name for The
a group of montmorillonite clay minerals including sodium, calcium and potassium
montmorillonites. These minerals, because of their chemistry and microstructure, are very
effective at absorbing water, swelling in the process and capable of absorbing several times their dry mass. Their structure is that of very thin flat crystalline sheets of negatively charged particles held together in stacks by positive cations in a layer of adsorbed water
(Milligan, 2000). Bentonite, once hydrated, remains strongly dispersed in water with a good
resistance to bleeding, i.e. settling out of suspension. Bentonite is very effective at filter cake
with as low formation, the creation of regions of permeability as of 10-‐10m/s in coarse-‐ grained soils possible. There are two classes of bentonite useful for construction, natural sodium bentonite and modified sodium bentonite. Wyoming bentonite is an example of a
natural sodium bentonite, with other regions around the world producing potassium and
calcium bentonite, which are less useful in construction. Calcium bentonite may be modified
These through ion exchange and transformed into sodium bentonite. ion-‐exchanged bentonites typically have the benefit of higher gel strength and viscosity than natural bentonite (Jefferis, 1992). Polymers are large, long-‐chain molecules formed by the linking together of smaller monomer molecules. Water-‐soluble polymers such as starch, guar gum, xanthan gum,
popular for ground engineering and have carboxymethyl cellulose or HPAM are becoming some advantages when used alone or mixed with bentonite. They are useful at much lower concentrations than bentonite, e.g. a 1kg/m3 polymer slurry may have similar viscosity as a
34
Chapter 2
Literature review
40 – 50kg/m3 bentonite slurry, and, since the long-‐chain polymer molecules effectively “encapsulate” the soil particles, polymers inhibit the swelling of active clay minerals (Jefferis, 2013). Other advantages offered by certain polymers include a greater resistance to salt and cement contamination and the ability to be broken down into less viscous liquids by the addition of simple chemicals like bleach (Jefferis, 1992), which may aid in waste disposal. Polymers may be added to bentonite slurries to improve their properties, and certain synergies may be formed with bentonite or other mineral clays. Polymers on their own or as foam mixtures may be introduced in relatively small proportions to ease the movement and transport of coarser soils, and are often used in this application in earth pressure balance tunnelling. However, polymers do not form filter cakes in the same manner as bentonite-‐based slurries, but rather rely on their high viscosity at low shear stress to suspend soil particles (Jefferis, 2013). Some filter cake may form with polymers, but it is typically a semi-‐permeable, membrane-‐like region of interwoven polymer chains enclosing soil particles, rather than a region of incrementally-‐built aligned clay platelets as is the case for bentonite (Majano et al., 1994). Slurries may exist in three states: dispersed, flocculated or gelled, as illustrated for the case of bentonite in Figure 2.16. When particles in the slurry carry an electrical charge, which depends greatly on the chemistry of the slurry, polar water molecules may become adsorbed onto the particle surface, forming a bound layer of water surrounding the particle. The repulsive force between the slurry particles are somewhat balanced by attractive Van der Waal’s forces, but not fully. So, the slurry particles are separated in suspension, in the dispersed state (Figure 2.16a). If the platy clay particles have different charges on the face and edge, they may be able to accumulate together in a clumped flocculated structure, as in Figure 2.16b. Certain additive may also facilitate this behaviour. These clumps, or flocs, settle out of suspension much more quickly than individual particles. Finally, certain types of slurry including bentonite-‐based slurry exhibit a time-‐dependent behaviour known as thixotropy. Here, the shear strength of the fluid depends on the shearing rate; if the fluid is left undisturbed, it “sets” to a certain gel-‐strength while it reverts to a viscous fluid when sheared (Milligan, 2000). The particle arrangement for this situation is shown in Figure 2.16c. The gel strength will depend on the particle concentration within the slurry, with higher concentrations of particles giving rise to higher gel strengths.
35
Chapter 2
Figure 6
Short-range forces between particles Literature review
Figure 7 -‐ Bentonite Structures bentonite(a) slurry (a) Dispersed (b) Flocculated (c) Gel Figure 2.16 slurry of structures: dispersed (b) flocculated (c) gel (Milligan, 2000) The study of these non-‐Newtonian fluids is complex, involving the study of soil mechanics, environmental engineering, chemistry and materials science. Since there is a such a wide range of slurry types, and slight differences in water chemistry may have such a large effect on their behaviour (for example magnesium ion concentrations of 50mg/L have been shown to cause problems in mixing of bentonite slurry (Jefferis, 1992)), this section is necessarily brief, however the properties presented and discussed should be sufficient to aid the reader with the understanding of the present research.
2.5.4 Pipe jacking lubricants Figure 8
Definition of viscosity
2.5.4.1 Reported mechanisms of action The term “lubricant” is perhaps a misnomer in pipe jacking, as “lubricants” are held to perform many functions (Norris, 1992; Schoesser et al., 2011), including: 1. Lubrication of the pipe string, so that the friction between the surface of the jacking pipes and the surrounding ground is reduced 2. Provision of buoyancy to the pipeline, so that the pipeline “floats” in a stable overcut 3. Stabilisation of the bore so that the surrounding soil does not transfer earth pressures to the entire circumference of the pipeline 4. Transport of soil or fractured rock particles, so that cuttings do not gather at the bottom of the overcut and cause the pipe string to stick. It has been shown that lubricants substantially reduce the force needed to drive a string of jacking pipes in the ground (Norris, 1992; Pellet-‐Beaucour and Kastner, 2002; Staheli, 2006;
36
Chapter 2
Literature review
Curran, 2010; Shou et al., 2010; Çetin et al., 2011), with reductions in skin frictional resistance of between 10% and 90% being attributed to the injection of lubricants. Four possible friction-‐reducing mechanisms are (Milligan, 2000; McGillivray, 2009; Schoesser et al., 2011): 1. The lubricant forms a lubricating boundary layer between the soil and the pipe. 2. The lubricant mixes with the soil to form a layer of material with a lower angle of friction. 3. The lubricant fills up the overcut, stabilising it and making the pipeline partially or fully buoyant. 4. The lubricant permeates into the soil until such time as a filter cake forms in the soil thereby reducing the effective stress on the pipeline.
These possible mechanisms will be examined at in detail in the following sections.
2.5.4.2 Film or boundary lubrication The ability of pipe jacking lubricants to provide “lubrication” during the driving of a pipe string through the ground is open to question. For comparison, the presence of water doesn’t “lubricate” soils, but instead the generation of excess pore pressures reduce the effective stress within the soil, and it is certainly the case that interface shearing resistance is not fundamentally affected by the presence of water (Dietz, 2000). It is felt that given the particulate nature of most soils, lubricants may not be able to form a sufficiently durable film or boundary to ease the movement of particles over each other.
2.5.4.3 Lubricant-‐soil mixtures The effect of bentonite slurry penetration into soil on the shear strength of the soil was investigated by Filz et al. (2004). Test results were presented for standard drained triaxial tests on coarse sand (D15 = 3.0mm) with water as a pore fluid and for similar sand samples, where the pore water has been replaced with 6% bentonite slurry through application of a pressure gradient. It was found that there was no significant difference in the strength of the consolidated coarse sand with water as pore fluid, or the same coarse sand with 6% bentonite slurry as the pore fluid, as shown in Figure 2.17.
37
slurry supply so that slurry was free to move into or out of the specimen in response to specimen volume change. As shown in Fig. 5, there is no significant difference between the results for the water-saturated specimens and the specimens penetrated with
Chapter 2
unit weight of the sand below th $ sw !unit weight of the sand when ter; $ s !unit weight of the slurry of groundwater; n!H s /H where Literature review the bottom of the trench; and &! sand. The inclination of the critica
' f !45°#
!
ta
Eqs. "2# follow the work of Nash and Amir-Tahmasseb "1965#; and that "1# different total unit weight above and below the groundwate surcharge pressure is included, a placed on the soil strength, i.e., ta the unit weight of the slurry. Fig. 7 shows a comparison b using Eqs. "2# and using Spence UTEXAS3 "Wright 1991# for a !20 m, H w !18 m, $ sw !20 kN q!0, and a range of values for $ the agreement between Eqs. "2# a 5. Shear coarse sand and swithout penFigure 2.17 -‐ Shear sFig. trength of cstrength oarse sof and with & wwith ithout lurry pslurry enetration (Filz t al., 2004) using UTEXAS3 Theecalculations etration gram to search for arbitrarily sha
JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENGINEERING ©
2.5.4.4 Buoyancy and contact conditions between the pipe and the soil Loganathan and Poulos (1998) base their elastic settlement analysis on the assertion that the pipe string settles to the bottom of the overcut gap, and hence the distance between the crown of the tunnel lining and the crown of the excavated space becomes twice the designed overcut. Norris (1992) and Marshall (1998) have observed pipelines both settling to the bottom of the overcut and undergoing buoyant uplift due to the action of water or slurry in the excavation. Here, the pipes were forced to the top of the overcut space and resistance to jacking was greatly reduced. Simple calculations carried out based on data provided by a concrete jacking pipe manufacturer (FP McCann, 2011) show that from nominal diameter (diamètre nominal) DN 900 upwards, concrete jacking pipes, if water-‐tight, are buoyant in water.
2.5.4.5 Filter cake formation As mentioned in Section 2.5.3, the ability of slurry to form a filter cake in the soil is of interest in pipe jacking and microtunnelling, as the formation of such a filter cake may be vital to one mechanism of skin friction reduction (§2.5.4.1). The formation of a filter cake depends on the relative grain sizes present in the soil and lubricating fluid, the pressure gradient between the pore water pressure in the soil and the injection pressure of the lubricating fluid, and time (Majano et al., 1994; Filz et al., 2004). It was found, for example, that at a pressure differential of 2kPa and after 4 hours, an unspecified bentonite slurry mix formed an effectively impermeable filter cake (Majano et al., 1994). Filz et al. (2004) indicate that a filter cake will form from bentonite-‐water slurry if the D15 size of the sand is 0.4mm or less and otherwise the bentonite-‐water slurry will penetrate into the soil. Schoesser et al. (2011) differentiate between an outer filter cake, inner filter cake and slurry penetration into the soil, as shown in Figure 2.18.
38
Chapter 2
Literature review
Figure 2.18 -‐ Outer filter cake (left), inner filter cake (centre) and slurry penetration (right) (Schoesser at al., 2011)
It is suggested that, for bentonite slurry, an outer filter cake forms when D10 is less than 1mm, an inner cake forms when D10 is between 1mm and 6mm and free penetration into the soil will occur when D10 is greater than 6mm.
2.5.5 Guidelines for lubricating pipe jacks and usage in the field 2.5.5.1 Field observations on the use of lubricants A search of the literature and discussions with pipe jacking contractors and drilling mud suppliers revealed that there were no explicit guidelines or commonly-‐applied rules as to what, when and how much to lubricate with. That lubricants should be mixed “to suit the ground” and applied “when necessary” was as far as could be established. The UK-‐based Pipe Jacking Association (2007) give the following guidance on the use of lubricants in pipe jacking: “Frictional forces on the pipeline may be reduced by applying a suitable lubricant, under a nominal pressure above that of the ground water pressure present.” Bentonite is the preferred choice of lubricant for most conditions within Ireland and the UK, normally in the form of CETCO Hydraul-‐EZ, a modified bentonite containing proprietary polymer additives. Some contractors mix bentonite with other polymer or surfactant products to combat swelling clays when necessary. Many contractor personnel interviewed had rules of thumb, for example to inject lubricant after every pipe or every other pipe at a set pressure above an assumed hydrostatic pressure. Others aimed to keep the annulus open, without stating exactly how this was achieved or checked for. Other researchers report similar difficulties in quantifying field lubricant use. Staheli (2006) indicated that the methods and lubrication schemes used by tunnelling contractors were based on experience in similar ground conditions, while Borghi (2006) stated:
39
Chapter 2
Literature review
“In spite of its importance, soil conditioning, or research thereon, is sometimes regarded with distrust and the ‘common sense and experience’ approach has tended to be preferred in the field. This attitude is perfectly reasonable for less demanding projects or in cases where only a low standard of performance is required. However, as the efficiency and control of the machine operation becomes critical or when the project conditions grow in complexity, a deeper understanding of the fundamental mechanics becomes crucial to complement the experience-driven approach to soil conditioning.”
Figure 2.19 – Automatic lubricant dosing system (Herrenknecht AG)
Recently, there have been attempts to manage and control the lubrication of pipe jacking operations. The Herrenknecht “Volume Controlled Bentonite System”, shown in Figure 2.19, is a new system that can automatically provide a dosage of a predetermined volume of lubricant along a pipe-‐jacked pipeline (Ulkan, 2013). Prior to the commencement of pipe jacking, the project designer prepares a lubricant dosage plan based on the geotechnical conditions expected, so that the optimal amount of lubricant is delivered. The plan is executed during the works, with all data recorded for later analysis. Meanwhile, the “CoJack” system allows the acquisition of jacking force data in real time and the processing of this data off site (Beckmann et al., 2007). In future, systems like these may make more data on lubrication practice in pipe jacking available for analysis.
2.5.5.2 Quantities of lubricant injected Pellet-‐Beacour and Kastner (2002) and the French Society for Trenchless Technology (2006) reported on the relationship between the volume of lubricant injected and the corresponding reduction in skin friction stress from a number of monitored drives. Volumes injected ranged from 25 litres per linear metre in marl and sandy-‐gravelly marl to 170 litres per linear metre in clean sand. In sandy soils, it is shown that volume of lubricant injected
40
Chapter 2
Literature review
and the percentage reduction in skin friction stress are closely related. It is also shown that continuous injection all along the pipe is the more effective strategy than reactive injections made when increases in jacking force were observed, as it may be too late, once the skin friction stress has increased, to have any effect with lubricant. A link between permeability of sandy soil, the volume of lubricant injected and the percentage reduction in skin friction resistance was also noted, showing that the quantity of lubricant injected must compensate for the more permeable sandy soils. Pellet-‐Beacour and Kastner (2002) present a graph of friction reduction percentage against the volume (litres) of lubricant injected per linear metre, which is reproduced in Figure 2.20, showing the beneficial effect of injecting large volumes of lubricant, particularly in sands. 88
A.-L. Pellet-Beaucour, R. Kastner ! Tunnelling and Underground Space Technology 17 (2002) 83!97
Fig. 8. Comparison between the reduction percentage of frictional stress and the volume of injected lubricant per linear metre.
Figure 2.20 -‐ Percentage reduction in skin friction stress against volume of injected lubricant per linear metre (ml = linear metre) (Pellet-‐Beaucour and Kastner, 2002)
great influence on the magnitude of friction stress A comparison between the different drives where addi reductions. This difference can be explained by a loss tional frictional stresses after an overnight stoppage For one site at Bouliac where the dominant soil type was “clean sand”, careful attention paid of lubricant, which may pass through more permeable were calculated, such as the analysis of variations of sandy and soils,allowed and byskin thefriction makingstress of a to remain to gravely lubrication at lubricant a very low value, on approximately injected amounts Montmorency 3 drive, soil!lubricant mixture on soil!pipe contact, which preshows the values of additional friction being reduced 0.5kPa, for the drive. Following this and other observations, the following factors were sents a larger friction coefficient than that of the lubriwhen the injected lubricant amount increases. Additioto inhave in that maintaining this low (French for toTrenchless cant. Datajudged collected sandyassisted soils show the volume nal stress frictional stressesSociety are seen decrease from 2.0 to of injected slurry and the reduction percentage of 0.6 kPa for increasing injection amounts from 26 to 200 Technology, 2006): l!lm Žsee Table 3.. frictional stress are closely linked. Of course, other These few examples show that an injection of slurry fieldwork monitoring will be necessary to quantify this relationship. in the annular space is effective for reducing additional • The provision of an adequate overcut, taking account of the possibility of swelling With regard to the injection process, it appears that friction after stoppages injections before work interrupclay and slosses ettlement ffects. control of there are some efficiency and epoor tions help to prevent the ground closure around the pipes. friction forces when injections are carried out discont• The addition of polymers and microbeads, which are small rigid spheres often added Therefore, the influence of lubrication was noted at inuously in reaction to critical increases of friction ŽPellet, twoslurry. levels: forces Žin example 1997t.o . In to Montmorency reduce drag in 3d. rilling fluids, bentonite fact, it is too late at this stage to achieve much benefit. • Continuous lubricant injection at several points along the pipeline, not just at the The lubricant injection distributed among several points ! Firstly, it reduces dynamic frictional stress in proportions varying from 45 to 90% Žsee Table 2.; and along the pipelinerear ensures a smore uniform distribution of the hield. of lubricant around the jacked pipes and a more effec! secondly, lubrication is effective in reducing additive reduction in frictional resistance. tional frictional stress due to work interruptions. In addition to its effect on dynamic frictional stress, lubrication is effective for reducing additional friction In fact, relationships between dynamic friction Ždur forces at restart, as the study on stoppages has shown. 41 ing jacking., additional friction resulting from stop-
Chapter 2 •
Literature review
A large injection volume (average injection rate of 5.6 times the volume of the overcut in the case of Bouliac, but in any case at least equal to the volume of the overcut).
•
Continuous jacking without long stoppages.
•
Careful measurement and monitoring of lubricant volume and pressure.
2.6 Modelling the skin friction component of jacking force 2.6.1 Approaches to modelling the skin friction stress in pipe jacking Many models have been proposed for modelling the skin friction component of jacking force in pipe jacking (McGillivray, 2009), ranging for simple empirical models to advanced analytical and numerical models. In common, almost all take account of the normal stress imposed by the ground onto the pipeline (§2.6.2). This will also depend on the stability of the ground, due to the overcut and arching, which may also be taken account of in modelling. A number of empirical models are reviewed, followed by the review of several theoretically based models of skin frictional resistance. For the theoretically based models, a distinction is made between coarse-‐grained soils and fine-‐grained soils, as models applicable to one may not be applicable to the other, while for empirical models the soil type is normally taken account of within the model itself.
2.6.2 Normal stress on underground pipelines 2.6.2.1 Overcut Normally, the pipe-‐jacking shield or microtunnelling machine is a slightly larger diameter than the product pipes, and the annular gap behind the shield or microtunnelling machine is termed the overcut or overbreak. The overcut reduces the normal stress on the pipeline. One of the following three conceptual mechanisms, or a combination, may influence the normal stress development following creation of the overcut: 1. The overcut may remain open, minimising the contact area and contact pressure between the pipe and the ground. 2. The overcut may yield, allowing for the development of arching mechanisms that reduce the normal effective stress on the pipe barrel. 3. The overcut may elastically unload onto the pipeline, or the swelling of clays may generate high stresses on the pipe barrel. These mechanisms are shown in Figure 2.21, while a probable physical scenario is also shown.
42
Chapter 2
Literature review
Figure 2.21 -‐ Conceptual mechanisms for overcut behaviour around a pipeline – open overcut (left), closed overcut due to collapse, swelling or elastic unloading (centre) and a possible physical scenario (right)
The overcut ratio (R) is the ratio of the shield diameter (DS) to the product pipe diameter (DP): 𝑅=
(𝐷! − 𝐷! ) 𝐷!
Equation 2-‐11
From a series of small scale tank tests using a model microtunnelling machine and coarse-‐ grained soil it was found that providing an overcut causes large reductions in jacking forces, reducing jacking forces on the model microtunnelling machine by a factor of up to 8 or 10 (Phelipot et al., 2003). Rogers and Yonan (1992) found that an overcut ratio of approximately 0.04 was optimal, with forces remaining low if the overcut ratio was increased beyond this. A factor of 3 to 5 reduction in frictional resistance between steel and two types of sand was observed using an overcut ratio of 0.0308.
2.6.2.2 Stability Once the overcut has been formed, its stability is of concern. The stability of the overcut is examined separately for the case of coarse-‐grained and fine-‐grained soils, but first the concept of arching in soils is introduced. 2.1.1.1.1
Arching
Since the effect of arching was first identified by French military engineers in the design of silos, where it was noted that the base of the silo supported only a fraction of the total mass above it and that the side walls carried a much greater load than expected, there has been much research into the phenomena (Lochaden, 2012). The active trapdoor problem, where movement takes places away from the soil mass, has long been used as a simple method to model the stress distribution occurring around tunnels. The classical active trapdoor problem defined by Terzaghi (1943) is reproduced in Figure 2.22.
43
s=c+utanq, Chapter 2
The unit weight of the soil is 'Y and the surface of the so q per unit of area. Literature review uniform surcharge The ratio between the 2
b 1
o'y
4=
n
t1h
I:
't
4
"-
.. I/)
:J
"
';=4
K=!
8
Figure 2.22 -‐ Sliding surfaces in Terzaghi's (a) trapdoor model (Terzaghi, 1943)
(d)
Once the overcut allows for enough resulting soil movement to initiate the arching
0
0
Z = tumdb
XI
2
mechanism, the vertical stresses acting on the crown of an underground pipeline may be calculated as follows:
"
2
𝜎! =
! 𝛾𝐵 /b
𝐾 ∙ tan 𝜙
I
II
.....
\ia!
Z
...... ......
\1
!
I':
I
𝜎! = 0.3𝛾 0.5𝐷! + 𝐷! 𝜎! 𝜋 𝜙 − 4 2
..... ......
Equation a I 2-‐12
! Klan Equation 2-‐13
6
b=c-Knf, Equation 2-‐14
8
-10)
Barla et al. (2006) – H’knecht AVN 1000
Limestone
558.3
Very high (>10)
Thomson (1993) – UK practice
Varies
-‐
2 to 3
Thomson (1993) – Australian practice
Limestone
-‐
1
Project/Reference
Rock type
Tullamore
Table 7.10 -‐ Observed face and skin friction stresses for the projects in Tullamore and Gilford (in italics) and values from literature
It is seen from the Table 7.10 that the face and skin friction resistance stresses measured at Tullamore and Gilford are broadly in line with expected values quoted in the literature. Of particular note is the discrepancy between the guidance on expected skin friction resistance stresses for pipe jacking in the UK and Australia given by Thomson (1993) and those observed in the field. It may be that this guidance needs updating. Also of note is the experience documented by Barla et al. (2006), where heavily jointed limestone of relatively low rock material strength but with strong micrite inclusions gave rise to very high face resistance stresses, which were high enough that one drive which was started using a microtunnelling machine with a pick-‐and-‐scraper head had to be abandoned with the machine being retrieved by open cut. The limestone also collapsed onto the pipeline generating very high skin frictional stresses. While the two case histories documented here were unremarkable, the information gained will be valuable to future microtunnelled pipe jacking works in hard rock. There remains an unfortunate lack of published case histories relating to this topic, but with the application of hard rock microtunnelling machines becoming more common it is hoped that further publications will arise in due course. Finally, neither in the two case histories described by the author nor in any of the case histories reported in the literature was the performance of the lubricants injected during the pipe jacking process measured or assessed. While it is probable that the Hydraul-‐EZ lubricant used in the projects at Tullamore and Gilford provided a contribution in keeping the pipeline buoyant and suspending unstable pieces of cut rock, it is impossible to verify this. Looking at the predictions made in Section 7.4.2, it is seen that the skin frictional
226
Chapter 7
Pipe jacking & microtunnelling in hard rock
resistance stresses recorded in the field were closest to the “full buoyancy” predictions than the “no buoyancy” predictions, which may indicate that the jacking pipes were buoyant in the lubricant to an extent. Given the cost and time involved in applying lubricants, it is suggested that some research be carried out in future to assess their effectiveness.
7.5 Summary The jacking forces recorded during two drives in rock have been back analysed. The total jacking forces were separated into face resistance and skin frictional resistance components. It is seen that skin friction resistance plays a minor role in the resistance to advance of pipe jacking pipes in rock compared to soft soil, while face resistance is higher. The skin friction resistance measured during the Tullamore project was close to the value of skin friction resistance predicted by assuming that the pipeline was fully buoyant in lubricant. This may indicate that the lubrication applied in Tullamore was effective in keeping the pipeline suspended within the overcut. For Gilford, the measured skin friction resistance stress lay between the predicted values for full and no buoyancy, perhaps indicating that the lubrication regime at Gilford was not as effective as that in Tullamore. It is seen that UCS serves as a useful basis for predicting the expected face resistance before a drive is started. An investigation of more advanced rock mass rating systems (RMR, RQD, QTBM, RSR) and their application to microtunnelling work is considered worthwhile, but was outside the scope of this research. It would also be worthwhile to undertake research to quantify the effect of lubricants in pipe jacking on hard rock. The availability of microtunnelling machines, such as the Herrenknecht AVN range fitted out with appropriate roller cutting wheels and cutting wheel changing facilities, has greatly improved the prospects for microtunnelling in hard rock. It is the author’s opinion that the availability of such machines should discourage attempts to excavate rock using traditional pick-‐and-‐scraper-‐based microtunnelling machines.
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8 Conclusion
8.1 Introduction As urban populations around the world grow and sensitivity to the disruption and environmental damage caused by traditional construction methods increases, trenchless technologies, including pipe jacking and microtunnelling, must continue to improve and adapt to meet the new challenges imposed. Ever more complex microtunnelling solutions are being conceived for projects such as long sea outfalls, seawater abstraction intakes and ambitious undersea oil and gas pipelines. These projects are increasingly demanding deeper, longer and more sharply curved pipe jacking drives, all factors which increase the jacking forces required. A thorough understanding of the mechanism of skin friction reduction due to the injection of lubricant slurries is important in solving the problems that will be faced. This thesis described a field, experimental and analytical study carried out to further the understanding of the mechanisms of skin friction reduction due to the injection of lubricant slurries into the overbreak during pipe jacking. From a review of the literature in the area, three main mechanisms have been determined which may give rise to the desired beneficial effect: 1. Boundary lubrication reducing the sliding resistance of the interface between the soil and the pipeline. 2. Reduction of the radial effective stress acting on the pipeline. 3. Rendering of the pipeline fully or partially buoyant in the lubricant slurry. The present research sought to clarify the extent of the contribution of each of these mechanisms to the measured skin friction stress reductions observed in the field through the analysis of field case history data and the physical modelling of the phenomena in the laboratory. In this chapter, the main findings and conclusions of the field and experimental work are summarised, the practical applications of the research are highlighted and areas where further work may be worthwhile are recommended.
228
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Conclusions
8.2 Conclusions 8.2.1 Field case histories Observations on eight lubricated microtunnelled pipelines in soft soil and two lubricated microtunnelled pipelines in hard rock have been presented. The jacking forces during the construction of these pipelines have been analysed, and the factors and operational parameters that may have affected the jacking forces have been described. For the soft soil drives, which were all lubricated, it was found that the skin friction resistances observed were generally low, and all were below 10kPa, which is a rule-‐of-‐ thumb established for estimating the skin frictional resistance in pipe jacking. Unfortunately it was not possible to differentiate between lubricated and unlubricated skin friction resistances, as records of lubrication procedures were not kept during the reported drives. Face resistance stresses were quite low for all cases, and were at the lower end of observed values reported by other researchers. It was concluded that the face resistance stress was dependent on the driving style of the microtunnelling machine operator, as well as soil conditions, with consistencies being established between operators of similar experience. The examination of field case histories in soft soils analysed by other authors indicated that, in ideal conditions, reductions of up to 93% in the skin frictional resistance were achievable with the injection of bentonite-‐based lubricants, with reductions up to and including 90% being commonplace. A method of estimating, within a broad range, the likely magnitude of the face resistance and the skin friction resistance stress during lubricated pipe jacking based on the type of soil expected during a drive was proposed, where the face resistance and skin frictional resistance from a number of projects carried out in Ireland and Northern Ireland were plotted against each other and grouped by a description of the dominant soil type during the drive. The data were subsequently plotted in a similar manner to cone penetration test results and this plot was compared with a well-‐known chart for the interpretation of soil behaviour from CPT results. The comparison suggested that the injection of bentonite-‐based lubricants during pipe jacking altered the behaviour of the soil around the pipeline, with the soil-‐bentonite mixture near the interface behaving like a sensitive, fine-‐grained soil. The jacking forces recorded during two drives in rock were back analysed. The total jacking forces were separated into face resistance and skin frictional resistance components, as for the soft soil sites. It was seen that skin friction resistance played only a minor role in the resistance to advance of the jacked pipeline in hard rock compared to soft soil, while face
229
Chapter 8
Conclusions
resistance was higher. It was shown that the uniaxial compressive strength of rock served as a useful basis for predicting the expected face resistance during a drive. The beneficial effects of lubrication in hard rock were shown in so far as the skin frictional resistance for one project was very close to that value predicted based on the assumption that the pipeline was buoyant in the lubricant within the bore. For the other project, the measured skin frictional resistance was between predictions made for the case of full pipeline buoyancy in lubricant and for the case of no lubrication, indicating that the lubricant had a significant beneficial effect in reducing the skin frictional resistance. It was concluded that Mechanism 3 quoted in Section 2.1, which was the rending of the pipeline partially or completely buoyant in the lubricant slurry, was the dominant mechanism of skin friction reduction due to the application of lubricants in hard rock pipe jacking. It was shown that modern microtunnelling machines, such as the Herrenknecht AVN range fitted with roller cutting wheels and cutting wheel changing facilities, were very effective in advancing pipe jacks through very hard rock. It was suggested, based on the examination of case history data presented by other authors, that the availability of such machines should discourage attempts to excavate even soft rock using traditional pick-‐and-‐scraper-‐based microtunnelling machines.
8.2.2 Physical modelling 8.2.2.1 Unpressurised lubricants In Chapter 4, the results of a series of unlubricated and lubricated direct shear interface tests on a number of coarse-‐grained and fine-‐grained soils and a rough concrete surface, similar in roughness to a typical concrete jacking pipe, were presented. It was found that for a fine to medium, silica sand, known as IGB sand, the roughness of the concrete interface used in testing was great enough that the full angle of internal shearing resistance of the sand was mobilised in the interface. In the case of the medium to coarse, calcareous sand, known as Banagher sand, the full angle of internal friction resistance of the sand was not mobilised because the particle size of the sand was larger. This was explained in the context of Uesugi and Kishida’s “critical roughness” model (1986a). The presence of an unpressurised layer of a bentonite-‐based lubricant was found to have a modest beneficial effect in reducing the shearing resistance between the test sands and the concrete interface. However, no measureable beneficial effect with regard to reducing the shearing resistance was observed in the case of fine-‐grained soils in either drained or
230
Chapter 8
Conclusions
undrained conditions. It was thought that this difference in behaviour was due to the particle size differences of the soils, in that the lubricant slurry could permeate the coarse-‐ grained soils and form a filter cake, whereas the void sizes present in the fine-‐grained soils did not allow this. The interface friction resistance between the concrete interface and the Banagher sand, placed at a relative density of 70%, was reduced by 6% when the concrete interface was coated in a 2mm layer of 8% Hydraul-‐EZ, while peak shear stresses were eliminated during shearing. In the case of IGB sand placed at a relative density of 70%, a 2mm layer of 8% Hydraul-‐EZ lowered the shearing resistance of the interface by 12.5%, while no reduction was seen with 4% Hydraul-‐EZ. Peak shear stresses were also eliminated during shearing of IGB sand against the rough concrete interface when 4% Hydraul-‐EZ was applied to the interface, but peak shear stresses are present in shearing following consolidation of the 8% Hydraul-‐EZ. For the coarse-‐grained soils under test, it was shown that the beneficial effect of shearing against a smoother interface outweighed the beneficial effect of coating a rough interface with unpressurised lubricant. In the case of a very smooth interface, which was a Perspex sheet, the mobilised angle of shearing resistance was around two thirds of the angle of internal resistance of each sand under test. In summary, the results obtained using unpressurised lubricants were not satisfactory in explaining the reductions in interface shearing resistance obtained in the field, in the case of either coarse-‐grained or fine-‐grained soils.
8.2.2.2 Pressurised lubricants In order to investigate the effects of pressurised lubricants on the shearing resistance of the pipe-‐soil interface, a novel triaxial testing procedure was developed that allowed for the injection of pipe jacking lubricants under pressure into the interface between a soil specimen and a concrete specimen, while shearing was ongoing between them under constant volume conditions. The development and commissioning of the apparatus and the development of the associated testing procedure were described in Chapter 5. The establishment of constant volume shearing conditions in the novel apparatus was considered important, as the soil around an advancing jacked pipeline was considered to be shearing at constant volume conditions, given the very large displacements involved. The results obtained from direct shear testing carried out under constant volume conditions, which were presented in Chapter 4, were used in calibrating the results of the initial,
231
Chapter 8
Conclusions
unlubricated, testing carried out with the novel apparatus. The application of the novel triaxial testing procedure to the simulation of lubrication in pipe jacking was then described, and the three lubrication strategies implemented were explained. These included static injection, where the lubricant was injected prior to shearing; injection at 4% axial strain, where the lubricant was injected during shearing at constant volume; and constant pressure injection, where a constant lubricant injection pressure was maintained throughout the shearing stage of the test. In Chapter 6, the results of testing carried out using the novel triaxial testing apparatus with a preformed failure plane were presented and the considerations taken into account when interpreting the results were explained. It was concluded that the novel apparatus accurately simulated the constant volume shearing of IGB sand in the interface with rough concrete, however the behaviour of the Banagher sand within the test procedure was not fully clarified. The magnitudes of reductions in the interface friction stress attributed to the injection of lubricants in the field were successfully replicated, and it was shown that the primary mechanism responsible for these reductions was Mechanism 2 given in Section 8.1, the reduction of the radial effective stress acting on the surface of the concrete interface. This was attributed the general increase in the pore fluid pressure near the interface due to the injection of either water or lubricant. Longer lasting interface friction reduction effects were seen when using viscous lubricant slurries than water, due to the lubricant slurries’ ability to clog the pores in the soil near the interface and hence allow for the retention of the lubricant pressure at the interface, even when the pressure in the lubricant system had been reduced. Polymer-‐ and bentonite-‐based lubricants, mixed at viscosities which resulted in Marsh funnel times of between 40 and 100 seconds, were both shown to be effective in clogging the pores in coarse-‐grained soils and facilitating the transfer of fluid pressures from the lubricant to the soil skeleton. The effect of differences in viscosity within this range was minor. The deviator stress ratio, which was the ratio of the lubricated deviator stress at 4% axial strain to the unlubricated deviator stress at that same strain, was the parameter by which the effectiveness of the lubricants during testing were compared. In static injection tests using IGB sand, reductions in the deviator stress ratio of 47% and 7% were observed when Hydraul-‐EZ and MX Polymer lubricants, respectively, were injected prior to shearing. In the case of Banagher sand, the reductions of the deviator stress ratio obtained during static
232
Chapter 8
Conclusions
injection tests were 94% and 78% for Hydraul-‐EZ and MX Polymer respectively. This showed that the lubrication process worked more effectively within the Banagher sand, and that the bentonite-‐based Hydraul-‐EZ was the more effective lubricant in the case of both sand types. For tests where lubricants were injected at 4% axial strain while the soil specimen was shearing under constant volume conditions, reductions in the deviator stress ratio of at least 94% were observed in all cases. Constant volume lubricant injection was only carried out with Banagher sand and in this case a 95% reduction in deviator stress ratio was achieved with Hydraul-‐EZ, while an 85% reduction was achieved with MX Polymer. Again, the Hydraul-‐EZ was more effective in reducing the interface friction stress in the case of Banagher sand, however both lubricant types produced results that were broadly in line with the reductions obtained in the field.
8.3 Summary This thesis has presented the results of a field and experimental study into the influence of lubricants in reducing the skin friction resistance in pipe jacking. The mechanisms of lubrication in soft soil microtunnelling have been shown to be both the boundary lubrication of the pipeline surface and the reduction of the radial effective stress acting on the pipeline due to the increase in the pore fluid pressure near the pipeline surface. It was shown that boundary lubrication was responsible for up to 20% of the obtainable reduction in interface shearing resistance due to lubrication in pipe jacking, with the remaining majority of the reduction being due to effective stress reduction at the pipe-‐soil interface. In hard rock microtunnelling, it was concluded that the buoyancy imparted to the pipeline due to the filling of the overcut with lubricant was the dominant mechanism for the reduction of the skin frictional resistance. The novel triaxial interface friction testing apparatus presented has provided a more comprehensive picture of the mechanisms of lubrication than previous research, involving constant volume shearing and the testing of two lubricants, one bentonite-‐based and the other polymer-‐based. The physical model developed simulates the behaviour of the soil near an advancing pipe jack in a more complete way than is possible using direct shear-‐based testing methods, allowing for the control and measurement of pore fluid pressures near the interface. A number of recommendations towards practice have been made, in particular that the monitoring of lubricant pressure during jacking should be made a routine part of the management of pipe jacking operations, and that lubricant strategies employed should be
233
Chapter 8
Conclusions
based on the maintenance of a constant lubricant pressure around the pipeline, rather than the injection of a prescribed quantity of lubricant. The traditional absence of documented procedures in practice for this very important aspect of pipe jacking is unfortunate, as lubrication, correctly applied, has the potential to reduce the frictional component of the jacking force by greater than 90%.
8.4 Recommendations for future research Although the research presented in this thesis has contributed to the knowledge of the effect of lubricants on interface friction in pipe jacking, more research is required to enable the full exploitation of the advances in pipe jacking and microtunnelling technology that have been made in recent decades. Opportunities for further investigation have been identified in the following areas: •
While the mechanisms of action of pipe jacking lubricants in coarse-‐grained soils have been clarified, the suitability of different lubricant types for different soil conditions has not been fully evaluated. Hence, further research is needed into matching the lubricant properties with the soil conditions.
•
Further investigations into the influence of lubricants on the behaviour of fine-‐ grained soils in the interface with jacking pipes are recommended.
•
The clogging effect of bentonite-‐based lubricant in particular is likely to be time dependent. It is recommended that consolidometer tests be carried out on bentonite lubricants and mixtures of lubricants and soils using low normal stresses to investigate time-‐dependent effects in pipe jacking lubrication.
•
Further field observation and back analysis is suggested, where lubricant pressure is continuously monitored at the injection port and at specific intervals along the pipeline, so the pressure dependency of the lubrication mechanisms may be validated. This would also allow a pressure profile to be built up along the pipeline so that the zone of influence of an individual lubricant port may be obtained.
•
Expansion of the triaxial interface shear testing programme may be worthwhile, using a larger test specimen incorporating a one or more extra pressure transducers to obtain further information on the behaviour of lubricant within the specimen.
•
The ability of a proprietary polymer product to clog the pores of a coarse-‐grained soil has been shown, however it is advised that further research look at the manufacture of an experimental polymer for use in physical modelling, such that its ingredients are known.
234
References Akram, M. & Bakar, M. Z. A. (2007). Correlation between Uniaxial Compressive Strength and Point Load Index for Salt-‐Range Rocks. Pakistan Journal of Engineering and Applied Sciences, 1, 1-‐8. Al-‐Mhaidib, A. I. (2006). Influence of shearing rate on interfacial friction between sand and steel. Engineering Journal of the University of Qatar, 19. Anagnostou, G. & Kovari, K. (1996). Face stability in slurry & EPB shield tunnelling. In: Mair, R. T., N. (ed.) Geotechnical aspects of underground construction in soft ground. London: Balkema. ASTM (2002). ASTM D5321-‐02. Standard Test Method for Determining the Coefficient of Soil and Geosynthetic or Geosynthetic and Geosynthetic Friction by the Direct Shear Method. ASTM International, PA, USA. ASTM (2007). ASTM D7012-‐07. Standard Test Method for Compressive Strength and Elastic Moduli of Intact Rock Core Specimens under Varying States of Stress and Temperatures. ASTM International, PA, USA. Baggeridge, P. (2010). Personal communication re. the construction of the Liffey Services Tunnel. Barla, M., Camusso, M. & Aiassa, S. (2006). Analysis of jacking forces during microtunnelling in limestone. Tunnelling and Underground Space Technology, 21, 668-‐683. Barton, N. (2000). TBM tunnelling in jointed and faulted rock. Rotterdam: A.A. Balkema. Bateman, G. (2008). Risk management at the Drogheda microtunnelling project. Trenchless Australia, June 2008. Beckmann, D., Stein, R., Fabri, T. & Uhlenbroch, A. (2007). CoJack -‐ A new statics method of computing and controlling pipe jacking. Tunneling and Underground Space Technology, 22, 587-‐599. Bieniawski, Z. T. (1989). Engineering rock mass classifications. New York: Wiley. Bishop, A. W. & Henkel, D. J. (1962). The measurement of soil properties in the triaxial test. London: Edward Arnold (Publishers) Ltd.
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Bolton, M. D. (1979). A Guide to Soil Mechanics. London: MacMillan. Bolton, M. D. (1986). The strength and dilatancy of sands. Géotechnique, 36, 65-‐78. Borghi, F. X. (2006). Soil conditioning for pipe-‐jacking and tunnelling. PhD Thesis, University of Cambridge. Bowden, F. P. & Tabor, D. (1958). The Friction and Lubrication of Solids. Oxford University Press. British Standards Institute (1990a). BS 1377-‐7:1990. Methods of test for soils for civil engineering purposes — Part 7: Shear strength tests (total stress). London: BSI. British Standards Institute (1990b). BS 1377-‐8:1990. Methods of test for soils for civil engineering purposes — Part 8: Shear strength tests (effective stress). London: BSI. British Standards Institute (2002). BS 5911-‐1:2002+A2:2010 -‐ Concrete pipes and ancillary concrete products. Specification for unreinforced and reinforced concrete pipes (including jacking pipes) and fittings with flexible joints. London: BSI. British Standards Institute (2004). BS EN ISO 14688-‐2:2004. Geotechnical investigation and testing -‐ Identification and classification of soil -‐ Part 2: Principles for a classification. London: BSI. British Standards Institute (2010). BS 1134:2010. Assessment of surface texture -‐ Guidance and general information. London: BSI. Broms, B. B. & Bennermark, H. (1967). Stability of clay at vertical openings. Journal of Soil Mechanics and Foundation Division, ASCE, 93, 71-‐95. Budhu, M. (2011). Soil Mechanics and Foundations, 3rd edition. Wiley. Çetin, D., Yıldırım, S. & Kılıç, G. (2011). A New Approach for Determination of Adhesion Coefficient in Pipejacking Method by Experimental and Numerical Analysis: An Example from Istanbul. Experimental Techniques, 37, 49-‐65. Chandler, R. J. (1966). The measurement of residual strength in triaxial compression. Géotechnique, 16, 181-‐186. Chapman, D. N. & Ichioka, Y. (1999). Prediction of jacking forces for microtunnelling operations. Tunnelling and Underground Space Technology, 14, 31-‐41. Colliat-‐Dangus, J. L., Desrues, J. & Foray, P. (1988). Triaxial testing of granular soil under elevated cell pressure. In: Donaghe, R. T., Chaney, R. C. & Silver, M. L. (eds.) Advanced Triaxial Testing of Soil and Rock, ASTM STP 977. Philadelphia, PA: ASTM. Craig, R. N. (1983). Pipe-‐jacking: a state-‐of-‐the-‐art review. London: CIRIA.
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Milligan, G. W. E. (2000). Lubrication and soil conditioning in tunnelling, pipe jacking and microtunnelling: a state-‐of-‐the-‐art review. London: Geotechnical Consulting Group. Milligan, G. W. E. & Marshall, M. A. (1998). The functions and effects of lubrication in pipe jacking. In: Negro Jr, A. & Ferreira, A. A. (eds.) ITA World Tunnel Congress 1998 -‐ Tunnels and Metropolises. Sao Paulo: Balkema, 749-‐744. Milligan, G. W. E. & Norris, P. (1996). Site-‐based research in pipe jacking‚ objectives, procedures and a case history. Tunnelling and Underground Space Technology, 11, Supplement 1, 3-‐24. Milligan, G. W. E. & Norris, P. (1999). Pipe-‐soil interaction during pipe jacking. Proceedings of the Institution of Civil Engineers -‐ Geotechnical Engineering, 27-‐44. Muir, A. J. (1971). Effect of shear strength parameters on slope stability analyses in the Winnipeg area. MSc thesis, University of Manitoba. Muir Wood, D. (1990). Soil behaviour and critical state soil mechanics. Cambridge: Cambridge University Press. Nicholls, H. (1929). The construction of a tunnel under the River Liffey. Transactions of the Institution of Engineers of Ireland, 60, 185-‐230. Norris, P. (1992). The behaviour of jacked concrete pipes during site installation. PhD thesis, University of Oxford. O'Kelly, B. C. & Naughton, P. J. (2008). On the interface shear resistance of a novel geogrid with in-‐plane drainage capability. Geotextiles and Geomembranes, 26, 357-‐362. O'Reilly, M. P. & New, B. M. (1982). Settlement above tunnels in the United Kingdom -‐ their magnitude and prediction. Tunnelling 1982 Symposium. London: Institution of Mining and Metallurgy, 173-‐181. O'Sullivan, D. (2010). Personal communication re. Bord Gáis Éireann gas pipelines under the River Liffey in Dublin. Orr, T. L. L. & Farrell, E. R. (1996). Geotechnical aspects of tunnelling in soft ground in Ireland. In: Mair, R. T., N. (ed.) Geotechnical aspects of underground construction in soft ground. London: Balkema, 301-‐305. Pellet-‐Beaucour, A.-‐L. & Kastner, R. (2002). Experimental & analytical study of friction forces during microtunneling operations. Tunnelling and Underground Space Technology, 17, 83-‐97. Phelipot, A., Dias, D. & Kastner, R. (2003). Influence of overcut and lubrication during microtunneling. ISTT International No-‐Dig 2003. Los Angeles. Pipe Jacking Association (2007). An introduction to pipe jacking and microtunnelling design. London: Pipe Jacking Association.
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Appendix A – Micro concrete mix design
Micro Concrete Mix Design Proserve Ltd.
Recommended Micro Concrete Mix 2:1 Sand:Cement ratio with the following mix proportions Cement Sand—Fine sharp washed. Zone 3 to 4 Water (mixing water:cement ratio 0.7, varies to suit sand grading)
556 kg/ m 1,112 kg/ m 387 kg/ m
Typical Strength Range 30—40 N/mm
80 Priory Road, Kenilworth, Warwickshire, CV8 1LQ, England UK: 01926 512222 Int: 00 44 1926 512222
[email protected] www.proserveltd.co.uk Fabriform
Recommended Sand Washed river or sea sand that is well graded within B.S. 822 Zone F. The sand grading line should fall inside the green zone on the graph below. 100 %
Percentage Passing (%)
90 % 80 % 70 % 60 % 50 % 40 % 30 % 20 % 10 % 0
0.075 mm
0.15 mm
0.3 mm
0.6mm
1.18 mm
2.36 mm
5 mm
Sieve Size (mm)
Flow Cone
The water:cement ratio varies with the exact granulometry of the sand used and should be determined on site using the Proserve Flow Cone. The water: cement ratio is adjusted to give a discharge time of 2733 seconds Proserve Flow Cone Volume 1.75 litres
Strength (kN/mm )
Free Water Bleed & Strength 41.4
Fabric Formwork
27.6
Conventional Formwork
13.8
0
14
28
The porous fabric allows the concrete to bleed, giving off its free water and thus reduces the pressures acting upon it. This allows the strength of the concrete to be higher than normal for a given concrete mix whilst retaining fluidity during placement B65
Days
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Appendix B – Hydraul-‐EZ technical data sheet
Hydraul-EZ Horizontal Directional Drilling Fluid
Description:
Hydraul-EZ is a high yield, 200-mesh sodium bentonite clay with a special dry polymer additive. It is designed to maintain borehole integrity in horizontally drilled boreholes. Hydraul-EZ is certified to NSF/ANSI Standard 60, Drinking Water Treatment Chemicals - Health Effects.
Recommended Use:
Hydraul-EZ is specially designed for conditions encountered in angle and horizontal drilling. It can be used for all types of freshwater mud rotary drilling and as a jacking lubricant.
Characteristics:
Mixing and Applications:
Mixes quickly. Requires less material due to low fluid loss properties. Concentrated for high yield. Forms a tight, thin filter cake in unstable formations. Maintains borehole integrity in horizontal and vertically drilled holes. Eliminates clay and shale swelling, bit balling and sticking problems. Mixing ratios are based on the use of fresh water. Water purity will affect bentonite performance. For best results, make-up water should be pre-treated with soda ash to a pH of 8.5-9.5. Hydraul-EZ should be added slowly through a jet/hopper mixer. Hydraul-EZ mixing ratios in pounds per 100 gallons of water: Normal conditions……………….20-30 lbs. Sand and gravel………………….30-40 lbs. Fluid loss control…………………40-60 lbs.
Bulk Density:
54 lbs/ft.3
Packaging:
50 lb. multiwall, water-resistant bags, 48 bags per pallet. All pallets are plastic-wrapped.
1500 West Shure Drive • Arlington Heights, IL 60004 • PH 847.392.5800 • FX 847.506.6150 The information and data contained herein are believed to be accurate and reliable. CETCO makes no warranty of any kind and accepts no responsibility for the results obtained through application of this information. REV 9/04
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Appendix C – MX Polymer technical datasheet Product data sheet MX LIQUID POLYMER SeaSafe for EPB.pages Issued: February 2009 Page 1 of 1
MX LIQUID for EPB in clean granular face Thixotropic, biodegradable, safe, ultra high performance polymer gel which conditions granular muck for proper support and flow Background and use MX LIQUID POLYMER for EPB meets the requirement for EPB face support in granular face. Very small additions of concentrate into the head produce a powerful gelling and lubricating effect. MX LIQUID POLYMER for EPB is effective in fresh and salt water. Formulation for most conditions would be 1 litre of MX LIQUID POLYMER per cubic metre of chamber volume.
Specific benefits True EPB pressure transmission Reduces torque Allows normal progress 100% biodegradable and non-toxic Waterproofing effect Highly effective in fresh or salt water Excellent stabilisation and carrying capacity in sands and gravels Excellent suspension properties when static or moving Safe for injection through the bulkhead
EPB this!
MX POLYMER was first developed to provide a superior replacement for bentonite slurry: Effective at very low concentrations No bulk mixing or transfer piping required Unlikely to compromise muck classification MSDS sheet on request.
Compatible with: All Morrison EPB products
Health and Safety
Packaging, form, availability
MX POLYMER has been specifically manufactured for safety in use, and is a very low-hazard formulation requiring no special labelling. CONSULT MSDS SHEET and assess safe use policy before use.
20 kg pails 200 kg drums Totes (IBC’s)
For further guidance, and specific recommendations and costings for your project, kindly contact your Supplier or: Mudtech Ltd - Morrison Mud Division, Wyburn House, 1 Crab Lane STAFFORD ST16 1SB United Kingdom
[email protected] www.mudtech.co.uk Tel: 44 1929 551 245 Fax: 44 1929 554 361
246
Appendix D – Triaxial testing procedure This Appendix sets out in detail the procedure used for the preparation of the specimens used in Chapter 5 “Interface friction reduction with pressurised lubricants – experimental apparatus and procedures”. Reference is made to Figure 5.13, which is reproduced below as Figure D.1.
Figure D.1 – Test set up for preformed failure plane triaxial compression tests
The procedure was as follows: 1. The tightly-‐fitting plastic tube was inserted into the pedestal of the triaxial cell and the concrete specimen with the preformed failure plane was set into place. 2. A piece of loose mesh, which was packaging material from supplies of laboratory filter paper, was loosely inserted into the tightly fitting plastic tube. This was to block the ingress of sand during the placing of the sand specimen later.
247
3. The 60ml syringe was filled with approximately 40ml of lubricant or water. Care was taken that the injection pressure line was free of foreign materials. With the valves on the injection pressure line opened, the lubricant or water was introduced to the circuit until it was just noticed escaping through the filter. Any excess lubricant was cleaned away using a paper towel. 4. The valve between the specimen and the injection pressure transducer was closed and the injection pressure transducer zeroed. The volume of lubricant or water remaining in the syringe was noted. 5. A membrane stretcher was used to place the thin latex membrane around the specimen and the membrane was secured around the pedestal using two rubber O-‐ rings. 6. A split-‐mould was set in place around the specimen and membrane. The top edge of the split mould was inspected for sharp edges and, if found, sharp edges were smoothed using sand paper so that the membrane wasn’t subsequently punctured. The loose end of the membrane was then stretched tightly over the outside of the split mould. 7. De-‐aired water was poured into the membrane until it was filled to the top. 8. A specimen of dry sand was prepared by weighing out the mass of dry sand required to fill the specimen, which had volume of 45cm3, to the predetermined relative density for the test. 9. The dry sand was placed in a beaker and sufficient water was added to more than cover it. The sand specimen was saturated and deaired in a vacuum dessicator. 10. The entire quantity of saturated sand in the beaker was transferred to the triaxial setup using a teaspoon. The sand was pluviated under water to avoid the entrainment of air during placement. Excess water was removed from the mould using a wash bottle as the sand displaced the water already in the mould. 11. A glass rod was used to compact the sand to a uniform relative density in three equal layers. The compactive effort required was established by experience with the particular sand under preparation. 12. Once placed, any loose sand sticking to the membrane or the split former was rinsed away using a wash bottle. 13. A saturated and de-‐aired sintered bronze porous disc was placed on top of the sand specimen. 14. At this point it was convenient to stretch two rubber O-‐rings over the top cap and let them run freely on the flexible drainage line. 15. The pore fluid circuit was evacuated to ensure that there was no air in the line, and the top cap was placed in contact with the porous disc.
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16. The membrane was gently stretched up over the top cap, and the two rubber O-‐rings were brought down over the membrane to secure it. The loose end of the membrane was neatly lapped down over the two rubber O-‐rings. 17. The split former was removed. Due to the fact the soil was saturated and compacted to a relatively high relative density, it was possible to leave the specimen temporarily free standing while the cell was assembled. 18. The cell was placed over the specimen, the air-‐bleed valve was opened and the cell was filled with water, before the air-‐bleed was closed so that the cell could be pressurised. 19. The test was ready to commence. The test procedures followed are described in detail in Chapter 5.
249
6S 6S 6S 6S 6S 6S 6S
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