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UBC Theses and Dissertations

Response of buried natural gas pipelines subjected to ground movement Weerasekara, Lalinda 2007

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R E S P O N S E O F B U R I E D N A T U R A L G A S PIPELINES S U B J E C T E D T O G R O U N D M O V E M E N T by L A L I N D A W E E R A S E K A R A B . S c , University of Peradeniya, Sri Lanka, 2002 A THESIS S U B M I T T E D IN P A R T I A L F U L F I L L M E N T OF T H E R E Q U I R E M E N T S FOR T H E D E G R E E O F M A S T E R O F A P P L I E D S C I E N C E in T H E F A C U L T Y O F G R A D U A T E S T U D I E S (Civi l Engineering) T H E U N I V E R S I T Y O F B R I T I S H C O L U M B I A February 2007 © Lalinda Weerasekara, 2007 ABSTRACT The performance of buried natural gas pipeline systems in areas subjected to ground displacements is an important engineering consideration since geotechnical hazards can be a major cause of damage to these utilities. With polyethylene (PE) pipes now becoming the industry standard for natural gas distribution systems, a detailed investigation into the interaction of these materials is needed to ensure that the response of both pipeline and soil components is properly understood during design. With this background, a detailed research program involving full-scale physical modeling has been undertaken at University of Brit ish Columbia (UBC) to study the effect of permanent ground displacements on buried P E piping. This thesis presents details on material properties, testing procedures and findings from laboratory full-scale tests performed to study the performance of number of buried pipe configurations applicable to field conditions. The results indicated that buried P E pipes subjected to relative soil movement in axial direction cannot be predicted using the simplified equations commonly used to capture the axial soil resistance in steel pipes. This is primarily due to complex pipe-soil interaction arising due to the flexibility and the non-linear material response of P E pipe, combined with dilation and arching effects from soil around the pipe. A new closed-form solution was derived to account the nonlinear material response of P E pipes. This solution provides the framework to obtain the i i response of the pipe (strain, force and the mobilized frictional length along the pipe) for a known relative displacement of the pipe. In branch pipe systems, the test results demonstrated that the physical location (branch pipe, trunk pipe or the tapping tee connection) of vulnerability would depend on relative pipe sizes, soil density, and boundary conditions. New analysis techniques were derived, and guidelines were proposed, to account the anchoring force offered by the tapping tee and the branch pipe in evaluating the performance of branch pipe configurations. In addition, technical details are discussed in relation to performance of the P V C protective sleeve and socket outlet tapping tees. These test results of straight and branch pipes would facilitate in developing a reliable database for calibrating/validating numerical and analytical models for potential future studies. i i i TABLE OF CONTENT A B S T R A C T i i L IST OF T A B L E S vi i i L IST OF F I G U R E S ix N O M E N C L A T U R E x i i i A C K N O W L E D G E M E N T xv C H A P T E R O N E I N T R O D U C T I O N 1 1.1 Background 1 1.2 Statement of problem for the present study 2 1.3 Objectives of the thesis 5 1.4 Scope of the thesis 6 1.5 Organization of the thesis 8 C H A P T E R T W O L I T E R A T U R E R E V I E W 10 2.1 Pipe-soi l interaction problem - Overview of analytical approaches 11 2.1.1 Closed-form solutions for pipe-soil interaction in axial and lateral soil loading 13 2.2 Research on buried steel pipelines subjected to ground movement 14 2.2.1 Soil load on pipelines from relative axial pipe movement 14 2.2.2 Soi l load on pipelines from relative lateral pipe movement 16 2.3. Field testing and monitoring of the response of buried pipelines 19 2.4. Research on buried Polyethylene pipelines subjected to ground movement 22 2.4.1 Findings from research at Cornell university 23 2.4.2 Findings from research at U B C 24 2.5 Summary of some of the pullout tests performed on pipes and anchor plates 27 2.6 Research needs identified from literature review 29 iv C H A P T E R T H R E E E X P E R I M E N T A L A S P E C T S 32 3.1 Test materials 32 3.1.1 Fraser River Sand backfil l 32 3.1.2. Properties of polyethylene pipes 35 3.1.2.1 The molecular structure of Polyethylene 36 3.1.2.2 Mechanical behavior of M D P E 36 3.1.2.3 The pipe diameter used in this study 40 3.2 Details of the apparatus/ test setup 41 3.2.1 Soil box 41 3.2.2. Actuators and hydraulic system 47 3.2.3 Instrumentation 50 3.2.3.1 Measurement of pullout resistance using load cells 50 3.2.3.2 Measurements of pipe displacement using string potentiometers 51 3.2.3.3 Measurement of pipe strain using strain gauges 51 3.2.4 Data acquisition system 52 3.3 Preparation of test specimens 53 3.3.1 Mounting of strain gauges on P E pipe 53 3.3.2 Placement of pipe in sand box 56 3.3.3 Placement of sand 57 3.3.3.1 Density of soil backfil l 58 3.4 Test program 63 3.5 Selection of test parameters 66 3.5.1 End boundary conditions for the branch pipe 66 3.5.2 Burial depth 67 3.5.3 Rate of ground displacement 68 3.5.4 The testing temperature 69 v CHAPTER FOUR AXIAL AND LATERAL PULLOUT TESTS ON STRAIGHT PIPES 70 4.1 Ax ia l pullout tests 71 4.1.1 Observed load-displacement-strain response for axial loading 72 4.1.2 Analysis of axial pullout resistance - Closed form solution 77 4.1.2.1 Pipe-soil interaction for an element level pipe length 78 4.1.2.2 Extension of element level interaction to model pipe response in a physical soil box test 79 4.1.3 Consideration in the selection of parameters to model for MDPE-sand interface at element level 89 4.1.4 Comparison of axial pullout resistance with other guidelines 97 4.1.5 Residual pullout resistance 99 4.1.6 Response of straight pipe subject to reverse loading 100 4.2 Lateral pipe pullout tests 101 4.2.1 Observed load-displacement-strain response for lateral loading 102 4.3 Summary 106 CHAPTER FIVE FULL-SCALE PULLOUT TESTS ON BRANCH PIPE SYSTEMS 107 5.1 60mm trunk pipes with 16mm branch pipes (Test nos. 60D16-U and 60D16-R) 108 5.2 114-mm trunk pipes with 16-mm branch pipes (Tests Nos. 114D16-U and 114D16-R) 113 5.3 114-mm trunk pipes with 42-mm branch pipes (Tests No . 114*42-*) 117 5.3.1 Induced rotation in the trunk pipe due to the tapping tee 123 5.4 Analysis of branch pipe configurations 127 5.4.1 Identification of components of branch pipe system for analysis 127 5.4.2 Anchoring resistance of the tapping tee (F,) 130 5.4.3 Resistance due to movement/anchoring of the branch pipe (Fb) 136 5.4.3.1 Resistance Fb in 16-mm branch pipes 136 5.4.3.2 Resistance Fb in 42-mm branch pipes 139 vi 5.5 Stress concentration due to tee and localized strains 141 5.6 Effect of rate of displacement on pullout resistance 144 5.7 Effect of soil density 145 5.8 Impact of the P V C protective sleeve 147 5.9 Summary 148 C H A P T E R SIX S U M M A R Y A N D C O N C L U S I O N S 151 6.1 Ax ia l and lateral pipe pullout tests 152 6.2 Pullout tests on branch pipe configurations 154 6.3 Future research requirements 157 6.3.1 Research requirements identified for axial soil loading 158 6.3.2 Research requirements identified for lateral soil loading 159 6.3.2 Research requirements identified for soil loading on piping networks 160 R E F E R E N C E S 162 A P P E N D I X A PIPE C O N N E C T I O N S : S A D D L E FUSION T A P P I N G T E E S 172 A P P E N D I X B A X I A L A N D L A T E R A L PIPE P U L L O U T TESTS R E S U L T S 175 A P P E N D I X C B R A N C H PIPE P U L L O U T TESTS R E S U L T S 197 A P P E N D I X D P H O T O G R A P H S , 235 vi i LIST OF TABLES Table 2.1 Summary of full-scale tests performed 27 Table 3.1 Details of P E pipes used in testing 40 Table 3.2 Details of the strain gauges 52 Table 3.3 Average densities measured in tests 61 Table 3.4 Test matrix 64 Table 5-1 Likely or observed failure modes and locations for the considered pipe configurations 126 vin LIST OF FIGURES Figure 2.1 Representation of pipe-soil interaction through (a) axial (b) lateral and (c) upward soil springs 1 Figure 2.2 Slide geometry depicted in current full-scale tests 2 Figure 2.3. Normalized axial pullout force and normalized displacement for 60-mm and 114-mm pipes in loose and dense sands ^ Figure 3.1 Grain size distribution of Fraser River Sand (Anderson, 2004) 3 Figure 3.2 Variation of the peak friction angle of F R S with confining stress (Karimian, 2006) 3 Figure 3.3 (a) Typical stress-strain behavior of Polyethylene and the effect of (b) temperature (c) strain rate, on the stress-strain behavior of Polyethylene (Stewart, et al, 1999) 3 Figure 3.4 Stress-strain characteristics derived from compression tests conducted on 60-mm and 114-mm P E pipes (Anderson, 2004) 3 Figure 3.5 The typical pipe configurations of (a) axial, (b) lateral and (c) branch pipe pullout test 4 Figure 3.6 Plan view of the soil box and the test setup 4 Figure 3.7 Side view of the soil box and the test setup 4 Figure 3.8 Two piece split- coupling used at the leading end of the 114 mm pipe 4 Figure 3.9 Formation of stress relief loop from lead wires of the strain gauge f Figure 3.10 Density measurements using nuclear densometer in (a) dense and (b) loose sand tests (Anderson, 2004) c Figure 3.11 Arrangement for unloading of sand from the sand box i Figure 3.12 Typical field installation of a distribution pipe system t Figure 4.1 Typical pipe configuration for axial pullout behavior of 60 mm and 114 mm bare pipes 72 Figure 4.2 Load-displacement characteristics for the 114 mm and 60 mm pipes 74 ix Figure 4.3 Displacement of the leading and trailing ends of the 114 mm and 60 mm pipes, (refer to Figure 4.1 for the location of SP1) 74 Figure 4.4 Strain distribution along the buried length of the 60-mm diameter pipe at different displacements of the leading end 75 Figure 4.5 Strain gauge readings for 60 mm straight pipe at a distance of 0.75 m from the leading end 76 Figure 4.6 Interaction behavior at the MDPE-sand interface 80 Figure 4.7 Mobi l ized elastic (x e) and plastic (x p) frictional lengths along the pipe 80 Figure 4.8 Stress-strain response for the 60mm P E pipe from uniaxial compression, axial pullout tests along with response modeled using hyperbolic model 83 Figure 4.9 Predicted and measured (from Test No . 60D) force-strain relationship at the leading end of the pipe in 60-mm pipe 86 Figure 4.10 Strain-frictional mobil ized length (x p) relationship derived from closed-form solution 87 Figure 4.11 Predicted and experimental force-displacement response in 60mm pipe 89 Figure 4.12 Comparison of predicted axial pullout resistance from A S C E (1984) for 60-mm and 114-mm pipes 92 Figure 4.13 Affected shearing zone in P E pipes, detected from the disturbances in the .strips of colored sand (Karimian, 2006) 96 Figure 4.14 Variation of K with pipe diameter (Karimian, 2006) 97 Figure 4.15 Prediction of the axial soil resistance using Danish Submarine Guidelines (1985) and McAl l is ter (2001).. 99 Figure 4.16 Predicted residual frictional resistance from loose soil properties (cp = 33° ,K=0.5) 100 Figure 4.17 Force-displacement relationship for the pulling (positive) and pushing (negative) directions 101 Figure 4.18 Placement of strain gauges in opposite ends of the bending plane, for lateral pullout tests 102 Figure 4.19 Test configuration for Test No . 114D-(2.5m) 103 Figure 4.20 Lateral load vs displacement relationship of Test No . 114D-(2.5m) and 114D-(1.5m) 104 x Figure 4.21 Bending moment distribution of along the length of the pipe in 114D-(2.5m) test, for different displacements of the pipe ends 105 Figure 5.1 Pipe configuration for Test No . 60D16-* and 114D16-* with strain gauge (SG) and string potentiometer (SP) locations in trunk pipe 108 Figure 5.2 Strain gauge readings (at crown and invert of the pipe) at a location of 0.75 m from the leading end of the pipe in 114L42-R test 110 Figure 5.3 String potentiometer readings at different locations of the pipe in 60D16-U I l l Figure 5.4 Strain distribution along the trunk pipe for different leading end displacements in 60D16-U test 112 Figure 5.5 Load -displacement characteristics for the 60D16-U and 60D16-R tests 112 Figure 5.6. Force-displacement characteristics of 114D16-R and 114D16-U tests 114 Figure 5.7. Strain distribution along the 114-mm pipe for different displacements of the leading end (Test No . 114D16-R) 114 Figure 5.8. Yie ld ing induced failure of the branch pipe at the tee, in Test N 0 . 114D16-R 116 Figure 5.9. Buckl ing induced failure of the branch pipe at the internal stiffener of the tee, in Test No . 114D16-U 116 Figure 5.10. Deformation of the branch pipe and movement of the P V C protective sleeve along the pipe in Test No . 114D16-U (Anderson, 2004) 117 Figure 5.11. Test configurations for 114*42-* tests showing strain gauge (SG) locations along the trunk pipe 118 Figure 5.12. Load-displacement characteristics of 114L42-U and 114L42-R tests 119 Figure 5.13. Load-displacement characteristics of 114D42-U and 114D42-R tests 119 Figure 5.14. Formation of plastic crack at the tee in 114D42-U test 121 Figure 5.15 (a) Punching failure at deep burial depths (b) Typical formation of passive and active soil wedges at shallow burial depths, when pipe moves laterally 122 Figure 5.16. Rotation of the trunk pipe in 114D42-R test 124 x i Figure 5.17. Torque induced in the trunk pipe due to the eccentricity (e) of the loading axes 124 Figure 5.18. Strain distribution along the trunk pipe for different displacements of the leading end, in 114D42-R test 125 Figure 5.19. Crit ical sections of the pipe for different total mobil ized friction lengths (xf) along the trunk pipe 129 Figure 5.20. Load-displacement characteristics of 114D and 114D42 (tee only) test... 131 Figure 5.21. Normalized lateral bearing capacity factors for different H/D ratios and friction angles ( A S C E , 1984) 133 Figure 5.22. Soil spring characteristics for the tee at a burial depth of 0.6 m and in dense sand 134 Figure 5.23. Measured and calculated (from soil spring behavior) pullout resistances of the 60D16 (tee only) test 135 Figure 5.24 Soil-spring characteristics for the unrestrained 16-mm pipe in dense sand 138 Figure 5.25. Measured force-displacement characteristics (at the leading end) from Test No . 60D16-U compared with those derived from soil-spring characteristics for tee and branch pipe 139 Figure 5.26 Branch pipe resistance of 42-mm pipe when restraint condition at the end restraint condition is mobilized 140 Figure 5.27. Local ized strain (in front of the tee) in the 60-mm trunk pipe due to presence of (a) tee, and (b) tee and 16 mm branch pipe (i.e. Test series 60D16-* with branch pipes) 142 Figure 5.28 Measured and calculated strain distributions along the trunk pipe from strain gauges and string potentiometer measurements (at a leading end displacement of 100 mm) in Test No. 60D16-U 144 Figure 5.29 Strains (at a distance of 0.75 m from leading end) in loose and dense sand when branch pipes are unrestrained 146 Figure 5.30 Strains (at a distance of 0.75 m from leading end) in loose and dense sand when branch pipes are restrained 147 x i i NOMENCLATURE A t Projected area of the tapping tee in a plane perpendicular to the direction of the movement. A p The cross sectional area of the pipe D Diameter of Pipe dso The average grain size of the soil, e Eccentricity in the loading axes in the tapping tee Fb Resistance due to movement/anchoring of the branch pipe. Fi Frictional resistance caused by the axial movement of the trunk pipe. F t Anchoring resistance of the tapping tee. / Interface Friction Angle Ratio H Depth to Pipe Springline K Lateral pressure coefficient ko Coefficient of Lateral Earth Pressure "at rest" L Embedded Length of Pipe N q Dimensionless horizontal bearing capacity factor P Transverse Horizontal Force 1 max Peak Ax ia l Pullout Resistance T x Ax ia l Pullout Resistance t Pipe Wal l Thickness u The relative displacement between pipe and soil u x The relative displacement at peak frictional resistance, T m a x xi i i Wp Weight of the pipe and its content x Ax ia l Displacement x e The mobil ized elastic frictional length along the pipe x p The mobil ized plastic frictional length along the pipe xmax Ax ia l Displacement at Peak Ax ia l Pullout Resistance yh Ax ia l Displacement at Peak Ax ia l Pullout Resistance 8 Interface Friction Angle y Unit Weight of Soi l cp Internal Angle of Friction of Soil a v Average normal soil pressure calculated from idealized pressure distribution xiv ACKNOWLEDGEMENT It is a pleasant aspect that now I have the opportunity to express my gratitude to all those who gave me unconditional support in completing this thesis. First and foremost, my sincere gratitude is extended to my supervisor, Dr. Dharma Wijewickreme for his unwavering guidance and encouragements towards the successful completion of this thesis. His detailed and constructive comments, understanding and enthusiasm paved the way for a pleasant research environment. Furthermore, I also greatly appreciate the support that Dr. Carlos Ventura has given in providing all the necessary software manuals and reviewing the thesis. The project was funded by the Terasen Gas Inc., B .C. M y warm thanks are due to Mr. A l len Mitchel l of Terasen Gas, Inc. for providing the necessary pipe materials and technical details promptly. His valuable comments and support was an immense help towards the completion of this thesis and in publishing technical papers on this area. The gratitude is also extended to the staff at Terasen Gas for their assistance and productive remarks in making this project a success. It is with great pleasure to recognize the valuable support given by the staff of the C iv i l Engineering workshop. I would like to express my deep and sincere gratitude to Mr . Doug Smith, B i l l Cheung, Doug Hudnik and Herald Schrempp of the C i v i l Engineering xv workshop for manufacturing the required testing components and for sharing their technical expertise. A lso a special recognition if given to Mr . John Wong and Scott Jackson of C iv i l Engineering electronic workshop for their support in instrumentation and strain gauging, and Mr . Max Nazar and Felix Yao of the structures laboratory of the Department of C i v i l Engineering for bearing the chaos during testing periods. Taking into consideration the immense difficulty of the work, my sincere gratitude is also extended to N ima Nabavi, Winson Cheng, James Lee, Kedar Deshpande and Al i reza Ahmadnia for their tireless efforts in test preparation and in emptying the sand box. I also owe my most sincere gratitude to fellow graduate student, Abdoul Hamid Karimian for all his input and advice during this project. xvi CHAPTER ONE INTRODUCTION 1.1. Background Since its introduction in late 1960's polyethylene (PE) pipes are becoming vastly popular in natural gas distribution industry. In North America, more than 90% of the natural gas distribution systems comprise of plastic pipes, of which 99% are P E pipes, (PIPA, 2001). Although H D P E (High Density Polyethylene) pipes are also used in practice; M D P E (Medium Density Polyethylene) pipes wi l l account for 2/3's of the usage in gas distribution industry. M D P E pipes have the added advantage of having higher flexibility and fracture toughness while having comparable long term strength and stiffness to that of H D P E (Stewart et al, 1999). P E pipes have become more popular due to its many advantages over more rigid steel pipes. Generally, the plastic pipes have a lower material cost, installation cost, maintenance cost, corrosion resistance, ease of processing, lower friction at the interface, lightweight and greater capacity to accommodate displacements than its counterpart, steel pipes (PIPA, 2001). 1 The performance of pipelines in areas prone to ground movement is a key concern for natural gas utility owners since the failures of the systems can cause property damage and even human losses, in addition to the invariably associated business disruption. Furthermore, the failures of the distribution lines occur in residential areas whereas, the transmission lines are so often placed in remote areas, away from locale or properties. Regardless of the smaller diameter in comparison to transmission lines, P E pipes are designed to carry significant gas pressures; (see Table 3.1 for maximum operating pressure for different pipe sizes) thus, in an event of a pipe failure, damage could become significant. Moreover, the prevention of such failures has been difficult and sometimes would become inefficient due to lack of current understanding of the response of buried pipes under installed conditions, ground movement, etc. 1.2 Statement of problem for the present study This focus of this thesis is on the likely pipe failures arising from ground movement. As a result of permanent ground movement, the failure in pipelines could occur due to stress concentrations (resulting brittle-like failures), yielding, buckling, etc. The ground movement can be either slow (few centimeters per year) or even can be a very fast liquefaction induced or mud f low type of movement, occurring at few meters per second. In the case of fast ground movements, very little can be done during a movement, unless proper precautionary measures have been taken in anticipation of such event. In contrast, a slow ground movement may provide adequate time to undertake necessary preventive measures, provided that the condition of the pipe is sufficiently known. 2 With the current technology such as GIS systems, the ground movements can be detected and mapped over time. However, even with precise ground movement data obtained from such methods, it is difficult to predict the condition of the buried pipe. One of the options to determine the pipe condition is to unearth the pipe, but this wi l l become a lengthy and a costly task. Therefore, methods should be developed to determine performance of the pipe from available ground displacement data. In view of developing pipe performance models, significant amount of research was conducted during last few decades, mainly focusing on high capacity steel transmission lines. Through laboratory research ranging from full-scale models to centrifuge tests, field testing and monitoring, analytical and numerical modeling, a relatively comprehensive knowledge base about pipe-soil interaction in steel pipes has been developed. Nevertheless, only very limited work has been undertaken to study the pipe-soil interaction with respect to buried flexible P E pipes. Due to the lack of alternative approaches, pipe-soil interaction models used in steel pipes are often considered for analyzing P E pipe configurations. These simplified models can not adequately represent rather complicated pipe-soil interaction in P E pipes. Considering the relatively smaller deformation stiffness than steel and time-dependent stress- strain response (viscoelastic and creep behavior) of P E as a pipe material, significant limitations would likely to occur when such methods developed for steel pipes are used to evaluate the response of P E pipes. 3 Joints and bends connecting the relatively small diameter branch pipes (laterals) to the trunk pipe are a common occurrence in natural gas distribution networks. A s seen from earlier tests performed by Anderson (2004), these joints may have significant impact on the integrity of the pipe system. Under certain situations, the localized stress concentration arising from ground movement due to the presence of these joints may become more critical than the friction induced stresses in the straight pipe sections. In addition, presence of many joints and bends accentuate that design basis used in transmission lines would not essentially adaptive to gas distribution networks, as such; the likely failure mechanisms and contributing factors to failure in distribution systems are sometimes significantly different to that of steel transmission lines. Clearly there is a need to undertake further research to understand the performance of buried P E pipelines subjected to ground movement. The outcome from such research would provide key input to developing the needed criteria or guidelines for design and performance evaluation of P E pipeline systems. With this background, a research program sponsored by Terasen Gas, Inc., B .C. has been undertaken at the University of British Columbia (UBC) to investigate the response of P E pipe systems under slow permanent ground movements. During the first stage of this work, a full-scale pipe-soil test facility was developed and several full-scale pipe pullout tests on selected diameters of P E pipes were performed (Anderson, 2004). In this work, initial studies on the response of straight pipes and branch pipe systems were conducted 4 for the cases where the ground movement is along the axis of the main pipe. It was noted that pipe-soil interaction behavior in P E pipes could become significantly different to that of steel pipes. It was also recognized that the performance of branch pipes may be further complicated by the type of pipe connection (e.g. saddle fusion tapping tee, see Appendix A) , under certain situations. In the current study additional detailed tests on branch pipes were performed and the contribution of such pipe systems was analyzed with the objective of understanding the performance of a system with multiple branch pipes. 1.3. Objectives of the thesis The main objective of this thesis is to conduct laboratory and analytical research leading to advance the knowledge with regard to the performance of P E pipelines subjected to ground movement. In turn, this understanding wi l l provide input to develop suitable guidelines and criteria to increase the level of confidence in evaluating the operational fitness of a given P E piping system. With this in mind, the following were identified as specific objectives: 1. To develop fundamental but comprehensive database for developing a numerical and analytical models to capture the basics of P E pipe-soil interaction. 2. To review some of the simplified equations stipulated in literature/guidelines in light of axial loading. 5 2. To identify the likely pipe failure mechanisms, critical locations, and key contributing factors to pipe failures when selected buried pipe configurations are subjected to specific relative ground movements, slide geometry and soil conditions. 3. To evaluate the impact of the presence of commonly used tee connections (mainly saddle-fusioned tapping tees - see Appendix A for "tee-connection" details) on a trunk pipe, in terms of additional resistance and induced localized strains for different trunk pipe sizes. 4. To technically assess the performance of commonly used tee connections (i.e. socket outlet type and butt outlet type) and other components such as P V C protective sleeves (see Appendix A ) under relative ground movements and review some of the current practices. 5. To examine the effect of relative sizes and boundary restraint condition associated with the performance of branch piping on the branch/trunk pipe system under relative ground movement. 1.4. Scope of the thesis As a result of very limited studies on pipe-soil interaction on P E pipe, variety of pipe-soil interaction aspects are needed to be addressed. However, on account of extremely time 6 consuming nature of the tests, the scope of study was limited to selected pipe configurations (pipe diameters and connections), soil and boundary conditions. In addition, fol lowing constraints on the scope are noted: 1. Only four pipe diameters (see Table 3.1) were selected for the testing. The selected pipe diameters were amongst the smallest pipe diameters in gas distribution industry. These diameters and accompanying tees/joints are common in field installations. 2. Generally in practice, burial depth varies between 0.3-1.5 m. Therefore, a burial depth of 0.6m was selected to replicate the field installations. 3. Cohesionless dry Fraser River Sand (FRS) was used in all the experiments, thus the effect of water content was not considered. Most of the tests were performed on dense backfi l l , considering the fact that pipes are more vulnerable in dense sand due to larger resistance than in loose sand. 4. The tests were conducted at a constant rate of displacement of 0.01 mm/sec (hence, the effect of rate dependency is not investigated), and ambient temperature around 20° C. This approach was to minimize the effects arising from strain rate and temperature dependant viscoelastic behavior of the pipe material. 7 The selection of these testing parameters wi l l be further discussed in section 3.2 of this thesis. 1 . 5. Organizat ion of the thesis Chapter 1: Introduction. The chapter presents the statement of the problem and the importance of this research. A brief description of the test is also provided together with the objectives and scope of the thesis. Chapter 2: Literature review. A description of previous laboratory full-scale pullout tests on straight pipes in axial and lateral directions, and pullout tests on branch pipe systems are given. Previous studies on steel pipes are also discussed under the topics of model tests, numerical modeling, analytical methods, field testing and monitoring, since this information is directly relevant for the understanding of the subject problem. Chapter 3: Experimental aspects. The chapter wi l l discuss the parameter selection for the tests with regard to burial depth, rate of displacement and boundary condition of the branch pipe and the resulting test matrix. The properties of soil and M D P E are discussed in details. Furthermore, this chapter wi l l include a description on the test setup, preparation, instrumentation and testing method. 8 Chapter 4: Ax ia l and lateral pullout tests on straight pipes. The chapter includes the test results (load, displacement and strain) for four axial pullout tests and two lateral pipe pullout tests. A closed-form solution was derived to account the nonlinear material behavior of the P E pipes. The axial pipe pullout test results wi l l be compared with some of guidelines available in practice. Chapter 5: Full-scale pullout tests on branch pipe systems. The test results are presented and discussed for three different pipe configurations. The potential and observed failure locations and major contributing factors for such pipe systems are discussed. With regards to pipe system analysis, the contribution of the tapping tee and the branch pipe to the pullout resistance are evaluated and recommendations are presented. Appendices The appendix A includes a description of commonly used pipe connections, and appendix B contains details of the test configurations and results of axial and lateral pipe pullout tests. The appendix C consists of test configurations and results for eight of the branch pipe pullout tests, whilst appendix D provides some photographs of different aspects of testing. 9 CHAPTER TWO LITERATURE REVIEW Numerous studies have been performed to understand response buried pipelines subjected to permanent ground movement. Thus far, most of this research has been focused on large diameter steel transmission lines, and the work has ranged from laboratory tests which spans from full-scale model tests to centrifuge tests, field testing, monitoring and numerical and analytical modeling. Although arising from steel pipelines, the findings from past work would stil l contribute to the understanding o f the pipe-soil interaction in P E pipes, as well as identifying areas where research would be necessary. This chapter presents a review of published information to describe our current understanding of the topic of pipe-soil interaction. Brief overview of analytical approach to modeling the problem is initially presented. Findings from research related to buried steel pipes subjected to ground movement are then presented. Limited information available on the performance of buried P E pipes subjected to ground movement is also included, thus forming the rationale for the present thesis. 10 2.1 Pipe-soil interaction problem - Overview of analytical approaches Modeling of pipe-soil interaction in the solution of real-life problem can be viewed at "element level" as well as at "regional level". The former pertains to the analysis of the response of specific buried pipe configurations. On the other hand, the latter corresponds to assessment of vulnerability of a pipeline network subjected to "regional" ground movements. The most popular (and also may be the easiest) method of "element level" analysis has been to represent the pipe-soil interaction by a set of discrete axial, lateral and upward nonlinear soil springs (as shown in Figure 2.1). Thus, the pipe-soil interaction can be conveniently analyzed by means of numerical techniques. These soil-spring models for longitudinal, lateral and uplift movements were obtained mainly from model tests performed on steel pipes (Audibert & Nyman, 1977, Trautmann and O'Rourke, 1985, Trautmann et al, 1985). Results of these model tests have been incorporated in: (a) Committee on Liquid and Fuel Lines Guidelines for Seismic Design of O i l and Gas Pipelines Systems, A S C E (1984) "Guidelines for Seismic design of O i l and Gas Pipeline systems", (b) American Lifeline All iance (2001) "Guidelines for the Design of Buried Steel Pipelines", and (c) PCRI (2004), "Guidelines for Seismic Design and Assessment of Natural Gas and Liquid Hydrocarbon Pipelines". These documents provide specific approaches to obtain these soil-springs, and some of these approaches are discussed in fol lowing sections. 11 Figure 2.1 Representation of pipe-soil interaction through (a) axial (b) lateral and (c) upward soil springs. Wijewickreme et al (2005) have described in details the approach for regional vulnerability assessment of a large natural gas transmission pipeline system (i.e. transmission system of Terasen Gas Inc., Vancouver, Canada). The method uses fragilities developed based on "element level" analysis, and then apply those fragilities to a system of pipelines via a GIS-based approach. In element level analysis, soil-springs developed as per above are typically employed in commercially available software, e.g. 3D general purpose programs as such as A B A Q U S , A N S Y S and N A S T R A N or in 2D pipeline analysis programs as P IPL IN , C E A S A R . During permanent ground movements, the response of given pipe configurations would depend on the slide geometry of the soil front, pipe orientation with respect to the soil movement, rate of ground movement, pipe's strength capacity and presence of joints, bends etc. 12 2.1.1 Closed-form solutions for pipe-soil interaction in axial and lateral soil loading From the theory of beams on elastic foundations (Hetenyi, 1946), closed-form solutions have been derived for longitudinal and lateral loading of pipes (O'Rourke & Nordberg, 1992, O'Rourke et al, 1995, Trigg and Rizkal la, 1994, Rajani et al, 1995, Chan & Wong, 2004). Using this theory, second and fourth order differential equations could be derived for longitudinal and lateral pipe movements and solved for known boundary conditions. Each method given in literature was different from each other due to different boundary conditions used in each method. From closed-form solutions for axial movement, it can be shown that inclusion of the elasto-plastic characteristic of soil and pipe material would prevent the design from becoming overly conservative (Trigg and Rizkal la, 1994, Rajani et al, 1995). For the lateral movement of a pipe, the closed-form solution is only val id for small ground displacements compared to the pipe diameter, thus could limits its applicability significantly. Chan & Wong (2004) incorporated the tensile force induced by bending into the solution. They also showed that the location of the maximum strain in the pipe moves along the pipe length with increasing strain. 13 2.2 Research on buried steel pipelines subjected to ground movement 2.2.1 Soil load on pipelines from relative axial pipe movement The recommended method for modeling soil load due to relative axial movement on straight pipes, buried in a cohesionless soil is a simple bilinear soil-spring with an axial soil resistance (TJ, which can be calculated from Equation 2-1 ( A S C E , 1984, A L A , 2001, PCRI , 2004). Formulation assumes an idealized soil pressure distribution around the pipe. The displacement at which the maximum force is mobilized (x m a x ) has been recommended on the. basis of experimental observations. Generally, in dense sand x m a x is suggested to be around 2-3 mm, whereas in loose sand this value would be around 5 mm ( A S C E , 1984). T = 7rDHLy(\ + k0)tanS [2_^ where, D = Diameter of the pipe H = Burial depth from ground surface to the springline of the pipe. L = Length of the pipe y = Unit weight of soil ko = Lateral earth pressure coefficient "at rest". 8 = Interface friction angle between pipe and soil. The interface friction angle is commonly represented as 14 S=f<p [2-2] where (p is the friction angle of the soil and / is a coefficient used to account for the reduction in soil friction at the pipe-soil interface. The accuracy of this soil spring model is limited, mainly due to difficulties in predicting the friction angle (8) and the lateral earth pressure coefficient (ko) at lower overburden stresses for different stages of axial loading. As a result, considerable discrepancies between predictions from Equation 2-1 and model tests have been reported (Pauline, et al 1998; Anderson 2004; Karimian, 2006). The relatively large friction angle measured at low overburden pressures is a contributing factor to this difference, e.g. from dry triaxial and direct shear test conducted at lower overburden pressures (15-50 kN/m ) the interface friction angle was found to be around 47.5° to 49.5° in very dense sand (see Figure 3.2) With the intention of reducing the interface friction between pipe and soil, Karimian (2006) also performed few axial pipe pullout tests with geosynthetic wrappings around the steel pipe. A significant reduction in friction was observed as a result of smaller friction angle between the two geosynthetics. As noted by Karimian (2006), the use of k 0 in Equation 2.1 does not adequately account for the normal stress from the soil mass around the pipe particularly during the occurrence of relative axial pipe movement. Based on pressure transducer readings 15 placed on the pipe, Kar imian (2006) found that for dense sand, the value of K can be much larger than suggested in literature. The soil dilation around the pipe would increase the soil pressure on the pipe, and the total increase is taken into account by a larger K value. In other words, when sheared, the lateral earth pressure wi l l be significantly different to that of "at rest" conditions due to soil dilation. Karimain (2006), suggested that the use of a " K " value, which is not necessary equal to ko but dependant on several other parameters such as, soil dilation, burial depth, etc. is more appropriate in the use of Equation 2.1 to calculate axial soil loads. As may be noted later, a lateral earth pressure coefficient (K) as suggested by Karimian (2006) is considered here. 2.2.2 Soil load on pipelines from relative lateral pipe movement Understanding the soil load during relative lateral pipe movement is another key component in modeling pipe-soil interaction. It is important to note that the lateral soil resistance is generally few magnitudes larger than the corresponding axial resistance. Examination of studies on lateral soil loading on pipes reveals a wide scatter among pipes having different pipe sizes, burial depths and backfil l soils. Under certain conditions, the differences in reported pullout resistances were more than 200% (Guo and Stolle, 2005). Hansen (1961) considered lateral movement of a rigid pile to formulate his analytical method for assumed shallow and deep failure mechanisms. Ovesen (1964) used pullout test results of vertical anchors plates to predict the lateral pullout resistance of pipes. In this method, a curved failure surface was assumed as opposed to the linear surface 16 assumed by Hansen (1961) and it was shown that Hansen (1961) method would overestimate the soil resistance. Up until Audibert and Nyman (1977) performed the first well-known lateral pullout tests on steel pipes, the pipe-soil interaction was predicted from model tests performed on vertical anchor plates. However, the stress distributions around the pipe and anchor plates are significantly different and the use of anchors to model pipes would result in overpredicting the soil loads on pipe. Neely et al (1973) and Das and Seeley (1975) also conducted series of tests on vertical anchor plates. Dicken and Leung (1979) conducted centrifuge tests on vertical anchor plates to examine the lateral pullout behavior. Due to limitations in predicting the pullout resistance from vertical anchor plates, Audibert and Nyman (1977) performed lateral pullout tests on 25, 60 and 114 mm pipes. They concluded that the passive resistance calculated from retaining wall analysis would not represent the ultimate soil resistance at shallow depths. Instead, Hansen's (1961) lateral bearing capacity factors (N q ) would represent the actual failure mechanism more accurately when the pipe moves laterally. However, after series of tests performed at Cornell University, Trautmann & O'Rourke (1985) concluded that their results are in agreement with Ovesen (1964). These tests results and Hansen (1961) have been included in developing A S C E (1984) and A L A (2001) guidelines. In these guidelines, it is recommended that the ultimate lateral soil resistance P u (per meter length) for cohesionless soil be calculated from Equation 2.3. P, = N ^ D [2-3] 17 where N q is the dimensionless horizontal bearing capacity factor and respective N q values can be obtained for different H/D values and soil friction angles from these guidelines. Using finite element analysis to model the pullout behavior of vertical anchor plates, Rowe and Davis (1982) showed that the ultimate soil resistance for lateral pullout wi l l depend on the burial depth regardless of whether separation would occur between back of the anchor plate and soil. Both cohesive and non-cohesive soils were analyzed using this method. Later, from test results of the centrifuge tests, Rizkal la (1992) pointed out that N q values are much higher than those recommended by Rowe and Davis (1982) and Hansen (1961). As described above, there is a considerable amount of variation in laboratory data (or predictions) was observed amongst many pullout tests performed under different pipe diameters, soils conditions and burial depths. Despite these variations in test results, Guo and Stolle (2005) showed from numerical analysis that these discrepancies in lateral pullout test results may be attributed to the scale effect, i.e. due to inconsideration of the different size of pipe diameters in dimensionless force. They also examined the effect of burial depth, overburden ratio and soil dilatancy on lateral pullout resistance. Karimian (2006) also conducted several model tests on 324 mm & 457 mm pipes and 2D numerical modeling to investigate the lateral pullout behavior. The findings supported the numerical predictions by Guo and Stolle (2005). In addition, few other numerical studies 18 have been performed by N g (1994), Zhou and Harvey (1996), Yang and Poorooshasb (1997), Guo and Popescu (2002), Popescue et al (2002). To date, very few tests have been performed on steel pipes to investigate the effect of rate of displacement on lateral loading. Hsu (1993) conducted a number of lateral pullout tests at several displacement rates and concluded that the rate effect could be represented by a power law relationship. Karimian (2006) has performed pullout tests on steel pipes buried in sand using rates of 2 mrn/s to 50 mm/s, and he found that no effect on displacement rates for the considered range of the loading rate. In addition to the axial and lateral movement of the pipe, the pipes would typically be subjected to relative movement in downward or upward directions, as a result of ground displacement. The downward movement is generally analyzed by bearing capacity theories for the soil. To analyze the uplift capacity of the pipe, few analytical, numerical and experimental methods have been proposed. Most notably, Trautmann et al, (1985) performed several upward pullout tests and presented a similar formulation to that of lateral movement. Although mentioned here for completion in terms of literature review, the uplift capacity of the pipe is not explicitly investigated in this thesis. 2.3. Field testing and monitoring of the response of buried pipelines As in general engineering problem, field pullout testing can be considered of significant merit since the tests are conducted on pipe essentially under "real-l i fe" conditions. Only 19 few field tests have been reported in the literature, and this primarily due to the fact that these tests are difficult in terms of logistics, in turn, becoming expensive proposition. Besides the difficulties and cost, the test conditions wi l l be much similar to the real life conditions. Audibert and Nyman (1977) performed a lateral pullout test on a steel pipe with a diameter of 230 mm. Rizkal la et al (1996) and Cappelletto et al (1998) have conducted field axial pullout tests on buried steel pipes with different backfi l l materials, ranging from cohesionless (granulite) soils to cohesive soils. From their studies, Cappelletto et al (1998) and Rizkal la et al (1996) have shown that A S C E (1984) recommendation for longitudinal pipe/soil interaction is overly conservative for cohesive soil backfil l. Another advantage in field testing is that they could be conducted over longer periods, without being subjected to space-time constraints that are more common in laboratory environments. A s such, field tests are preferable when the focus is towards understanding the effects of time performances. For example, Cappelletto et al (1998) noted that soil aging could also increase the dilatancy, and in turn the pullout resistance. The effect of aging wi l l be particularly important in cohesive soils where the peak resistance could increase by a factor of 1.8-2.0. Cappelletto et al (1998) also suggested that, in very slow ground movements, the expected strains at large displacements are closer to the peak value rather than the residual since new bonds are formed and cementation would occur at slow displacement rates. 20 Under certain circumstances, continuous field monitoring of critical gas pipeline sections is indispensable, even i f it becomes tedious. Bruschi et al (1996) and Bughi et al (1996) reported details of such field monitoring and these test results were subsequently used in numerical modeling of the actual pipe-soil interaction. Nevertheless, numerical modeling would become complicated due to lack of knowledge in representing the slide front and due to difficulties in determining the actual pipe-soil interaction taking placing because of varying soil conditions (Bruschi et al, 1996). Bughi et al (1996) also reported that soil constitutive law, boundary conditions and the assumptions in initial soil state can significantly affect the results of a numerical modeling. Furthermore, in remote areas or i f the ground movement is spread over a wide region, the continuous monitoring of the pipeline would become costly and unfeasible. In order to use any pipe-soil interaction model, it is essential to identify the sliding geometry of the soil mass as well . A typical landslide would consist of a separation area at the top and a soil accumulation area at the bottom of the slide. If the slide is along the pipe axis, in the soil separation zone, the sliding soil block w i l l be separated from the stable ground, thus creating tensile stresses in the pipe (see Figure 2.2). However, at the base of the landslide, due to soil accumulation, the pipe wi l l be under compression (Trigg & Rizkal la, 1994). The slide geometry replicated in the full-scale tests is similar to that of field landslide scenario depicted in Figure 2.2. Hence, the selected slide geometry limits the applicability of these test results to more general slope failure. However, as noted by Bughi et al (1996) generally in slow ground movements, most of the slip failures would occur at a depth below 6 m. Since the pipes are generally buried at a depth of 0.3-1.5 m, 21 the soil around the pipeline would be stable unlike in rapid flows or liquefaction induced flows (where the soil is unstable around the pipe). Thus, in most of these slow ground movements, the slide geometry can be represented by two soil masses separating from each other. The test results wi l l become valid only for the stable soil mass, i f one soil mass is unstable in the immediate vicinity of the pipe. Figure 2.2 Slide geometry depicted in current full-scale tests. 2.4. Research on buried Polyethylene pipelines subjected to ground movement So far, only a very few studies have been performed to evaluate the performance of P E pipelines subjected to ground movements. Unlike for the case of steel pipes, many factors primarily arising due to the pipe material type could affect the pipe-soil interaction response in P E pipes, i.e. the viscoelastic properties of the pipe (temperature and strain 22 rate dependency) and diametric changes arising due to relatively low deformation stiffness and associated Poisson's effect compared to steel pipes. To the best of author's knowledge, only two research institutions (Cornell University, Ithaca, N Y and University o f Brit ish Columbia, Vancouver, B.C.) have undertaken research on soil-structure interaction aspects of P E pipes. As such, the following two sections summarize the research findings of these two universities. 2.4.1 Findings f rom research at Corne l l university Stewart et al. (1999) reported the observation from a laboratory research testing program on P E pipes at Cornell University, in which, his investigation was focused mainly on the thermal effects and the behavior of P E pipes under temperature-induced cyclic loading. The axial pullout resistance of buried 150-mm diameter H D P E pipes was measured in a temperature-controlled test compartment at different temperatures. The test results indicated that temperature can have very significant impact on the frictional resistance due to diametric changes in the pipe. A t lower temperatures, the reduction in pipe diameter would reduce the normal stress on the pipe as soil arches around the pipe, in turn, lowering the frictional resistance at the pipe-soil interface. The results also revealed that the thermal effects on pullout resistance would depend on the pipe diameter. In large diameter pipes, the change in resistance would be large due to proportional diametric reduction or increase. Moreover, the cyclic tests (repeated pull ing and pushing tests) on buried P E pipes showed that soil resistance degrades considerably with repeated loading. 23 There was very small recovery of shear resistance due to aging after cycl ing; it was, however, noted that some resistance could be recovered through surface vibration. As a part of research on pipe-soil interaction, O'Rourke et al. (1990) performed over 450 direct shear tests on different sands to determine the interface friction angle for various plastics. The tests were performed at overburden pressures of 3.5-35 kPa. A s observed for M D P E , the (j)/8 varied only between 0.55-0.65 despite the different § and 5 values obtained for different sands and overburden pressures. Higher values of o)/8 was observed in sands with sub-angular particles. Interestingly, in the work reported by O'Rourke et al. (1990), there was no significant dilation or volumetric strains observed at the PE-soi l interface, even at large relative densities. This was assumed to be due to the skidding of the particles along the relatively hard interface, instead of roll ing as in softer surfaces. In softer surfaces such as P V C , dilation was observed at the interface. It was also seen that only 2-3 mm of displacement is required to achieve the residual frictional stresses. 2.4.2 Findings f rom research at U B C The studies performed on P E pipes by Anderson (2004), comprised of a number of axial pullout tests on straight pipes (i.e. diameters 60 mm and 114 mm) and four pullout tests on branch pipe systems. The results from the axial pullout tests of the straight pipes (60 mm and 114 mm diameters) are shown in Figure 2.3. It was observed that the peak axial 24 load would develop at axial displacements in the order of 20-mm. This was significant contrast to steel pipes, where the peak axial resistance would develop at relatively smaller displacement (i.e. 2-3 mm). In steel pipes, the pipe acts as a rigid member, thus x m a x (the displacement at which peak resistance is mobilized) is not affected by the elastic/plastic elongation of the pipe. Anderson et al, (2004) showed that axial force predicted from A S C E (1985) and A L A (2001) would overpredict the resistance in loose sand and underpredict in dense sand for the considered soil and interaction properties. However, it should be noted that the predicted magnitude of the peak soil resistance would depend on the accuracy of the parameters selected in Equation 2.1. In general, the predictions by Anderson (2004) were made with very limited understanding about the normal soil stress on pipe, or coefficient of lateral earth pressure coefficient (K). Anderson (2004) also reported the results from tests on branch pipe configurations. In these, 114-mm and 60-mm diameter trunk pipes, with two 16-mm branch pipes separated from a distance of 1.5 m from each other, were tested. In all tests, the pipes were instrumented with strain gauges at selected locations along the trunk pipe. The branch pipes were connected to the trunk pipe via saddle fusion tapping tee connections. The tests revealed large strains closer to the tapping tee. From these observations, it was concluded that in smaller diameter trunk pipes (60 mm dia), the trunk pipe is more vulnerable than the branch pipe. In contrast, the branch pipes would be at risk when used with large diameter trunk pipes (114 mm). However, in all four branch pipe tests, the end 25 of the branch pipes were unrestrained (not fixed at a given length), thus branch pipes could move freely with increasing trunk pipe displacements (see pipe configuration in Figure 5.1 for more details). Anderson et al (2003) also noted that the branch contribution (difference in pullout resistance between straight pipe and the branch pipe system) was independent of the diameter of the trunk pipe. 3.0 0.0 115aD 60aD—115aL 60aL 2.0 4.0 6.0 Normalized SP Displacement, x/D 8.0 Figure 2.3. Normalized axial pullout force and normalized displacement for 60-mm and 114-mm pipes in loose and dense sands (Anderson, et al. 2004) 26 2.5 Summary of some of the pullout tests performed on pipes and anchor plates It would be of interest to examine the laboratory tests performed on pipes and anchor plates, and following table summarizes selected full-scale testing that contributed to the understanding of the pipe-soil interaction. Table 2.1 Summary of full-scale tests performed (a) Tests on vertical anchor plates Reference D and (L /D ratio) H/D ratio Type of tests Soil properties ( 0 and density) Ovesen(1972) 15-150 (plain strain) 1-10 Lateral 15.2-16.7 k N / m J O = 32-43 Neely(1973) 51(5) 0.5-4.5 Lateral 15.8 k N / m 3 $ = 35-39 Das & Seeley (1975) 51(5) 0.5-4.5 Lateral 15.8 k N / m J 0 = 34 Dick in & Leung (1983) 25-152 (5) 0.5-6.5 Lateral (centrifuge) 16 .4kN/m J 0 = 4 4 27 (b) Tests on steel pipes Reference Pipe diameters H/D ratio Type of tests Soi l properties Audibert & Nyman (1977) 25-60-114 mm and 230 mm (field test) 1-24 Lateral 15.7-17.3 kN /m 3 O = 33-40 Trautmann & 0'Rourke(1985) 102, 324 mm & 1.2 m long 1.5-22 Lateral 14.8-17.7 kN /m 3 0 = 31-44 Trautmann, et al (1985) 102 & 324 mm 1.2 m long 1.5-11 Upli f t 14.8-17.7 k N / m J 0 = 31-44 Hsu (1996) Hsuet al (2001) 38.1-304.8 mm and 1.22 m long 1-12.5 Lateral & oblique 15 .2kN/m J 0=33-42 Paulin et al (1998) 328 mm N R Longitudinal and lateral N R Karimian (2006) 324 & 457 mm 2.5 Longitudinal and lateral 15.3-16.1 k N / m j 0 = 31-44 (c) Tests on P E pipes Reference Pipe diameters H/D ratio Type of tests Soi l properties Stewart, et al(1999) 150 mm 7.6 Longitudinal 17 .0kN/m j Anderson, (2004) 60 mm and 114 mm 5.5 and 10 Longitudinal and branch tests 15.7-16.3 0 = 32-38 * N R : Not reported 28 2.6 Research needs identified from literature review From literature review, it was recognized that only few tests have been performed on small diameter P E pipes to investigate the pipe-soil interaction behavior. Several areas of research have been identified as a result of lacking research in the topical area. In view of analyzing a pipe network subjected to slow ground movement, it was recognized that in addition to testing of complex pipe networks, individual pipe components such as axial and lateral movement of the straight pipe sections should also be tested. 1. In axial pullout situations, many factors could affect the pipe-soil interaction in comparison to steel pipes and these discrepancies are reflected in actual model test results and in corresponding predictions made using simplified equations. For the basic understanding of the pipe performance, a thorough analysis is required with the fol lowing factors investigated under carefully controlled test conditions. (a) The effect of the soil dilation on the lateral earth pressure coefficient (K) is essential in understanding the response of a pipe under axial soil displacement. Also the possible diametric reduction and resulting arching of the soil around the soil, when the pipe is subjected to axial strain. (b) The impact of the nonlinear (visoelastic) stress-strain behavior of the P E pipe needs to be examined. This may include investigation of the impact of rate of displacement and temperature on interaction behavior. 29 2. Very little is known about the lateral movement P E pipes. Model tests and analysis are required to determine whether the existing guidelines for steel pipes would sufficiently capture the behavior of P E pipes. 3. In branch pipe systems, in addition to investigating the effect of pipe diameters, soil condition and burial depth of the pipe systems, the following key factors should also be examined. (a) As the first step in analyzing a gas distribution system, identifying the key contributors to failure and critical sections of the pipe, for a given soil, pipe and boundary conditions is indispensable. (b) The impact and performance of different types of connections (e.g. butt tee connection, socket outlet and butt outlet tapping tees) should be examined, since in some situations, the performance of the tee may govern the overall performance of the pipe systems. This includes estimation of localized strains and anchoring forces induced in the trunk pipe, due to anchoring of the tee connection. (c) The impact of the end boundary condition and the corresponding resistance offered by the branch pipe should be evaluated for different pipe systems. 30 Considering the above factors, it was identified that the effect of soil dilation in axial pullout directions requires more extensive studies, and the limited work performed during this work wi l l not be sufficient for a quantitative assessment of this impact. The limited number of tests was performed in the lateral direction to obtain the necessary test parameters for numerical modeling. 31 CHAPTER THREE EXPERIMENTAL ASPECTS Considering the objectives and scope of the study, a test program was developed to investigate the pipe-soil interaction in P E pipes. The research is to examine the performance of branch pipe systems and fundamentals of pipe-soil interaction in straight pipes. This chapter presents the experimental aspects, including the following: (a) materials used in testing (sand and pipe materials), (b) their properties (friction angle, density, stress-strain behavior, etc), (c) details with regard to selection of testing parameters, i.e. the rates of displacement, end boundary conditions, burial depths, etc, and (d) testing methodology in relation to test preparation (soil handling, strain gauging, etc) and instrumentation (load, displacement, strain gauge measurements). 3.1 Test materials 3.1.1 Fraser River Sand backfill In all tests, dry Fraser River Sand (FRS) was used as the soil backfi l l , which is a material that has been extensively used in laboratory research at U B C over the past 10 years. The moisture content of the samples of F R S was measured to be less than 1%. The soil was kept at this low moisture content since: (a) maintaining a uniform water content 32 throughout the sand box would become difficult under moist condition; and (b) analysis of pipe-soil interaction would also become more complex due to shear induced pore water pressures. The mineral composition of the F R S was found to be 40% quartz, 11% feldspar, 45% unaltered rock fragments and 4% other mineral and characterized as fine to medium sand (Vaid and Thomas, 1995). The grain size distribution shows a minimum and an average grain size o f 0.074 mm and 0.23 mm, respectively (see Figure 3.1). The coefficient of uniformity (C u ) was found to be 1.5 with a specific gravity (G s) of 2.70 (Anderson, 2004). It is further assumed that the grain size distribution was not altered significantly due to repeated compaction of sand. This assumption has already been shown reasonable based on the comparison of grain size distributions (were almost similar) obtained before and after testing (Karimian, 2006). The minimum and maximum void ratios of F R S are 0.62 and 0.94, respectively. Furthermore, the grains are angular to sub-rounded in shape (Wijewickreme et al., 2005). The constant volume friction angle has noted to be 32-34° (Uthayakumar, 1996, Sivathayalan, 2000). 33 10 1 : 0.1: 0.01 0.001 Grain Size (mm) Figure 3.1 Grain size distribution of Fraser River Sand (Anderson, 2004) The fr ict ion angle of sand The strength deformation properties for Fraser River Sand were available from Karimian (2006). In his study, several dry triaxial tests were performed to find the friction angle of FRS at low overburden stresses. Figure 3.2 shows the variation of friction angle for two soil densities, 1665 and 1575 kg/m 3 at different overburden pressures. From these test results, the variation of friction angle for a soil density of 1610 kg/m 3 (i.e. corresponding to the average density used in dense test specimens of this program) could be interpolated. Based on such extrapolation, the friction angle of sand for a soil density of 34 1610 kg/m and typical overburden pressures of 10 kPa at the springline of the pipe was determined to be about 46.5°. 5s — 50 49 48 4? 46 45 44 43 42 •3  + 10 i i i i i i i i i Density. 1665 kg /m 3 i i i ^,*,******,B-4-»-riz—"~ ^ i i i i i i i i i i i i i i i ! liitrapolated VJV.. T t , 1.,.. __ ^ 3 > t — ' 1 Density, 1575 kg / in" ^ ^ * r * ^ ^ - - » : J i i i 1 1 1 1 20 30 40 Confining Stress (kPa) 50 Figure 3.2. Variation of the peak friction angle of FRS with confining stress (Karimian, 2006) 3.1.2. Propert ies of polyethylene pipes Unlike in rigid steel pipes, the properties of PE pipes could have significant effect on the pipe-soil interaction behavior. This aspect will be further discussed in the Chapter 4. The following will focus on determining the proper stress-strain behavior and interface friction of PE pipes. 35 3.1.2.1 The molecular structure of Polyethylene Polyethylene is produced by polymerization of ethylene molecules. The molecular structure of polyethylene consists of crystalline and amorphous formations. Crystalline long chains of molecules form the backbone of the material (Powel, 1983). The amorphous short branches extending from these molecules form the cross-links with the adjoining crystalline molecules. The short and long-term performance depends on the behavior of these amorphous and crystalline structures under loading. The different properties of high, medium and low density P E can be explained from their respective molecular structure. The following discussion wi l l focus only on the properties of M D P E , which is widely used in the gas distribution industry. 3.1.2.2 Mechanical behavior of MDPE As documented by many, the mechanical properties of M D P E depend on time (stress-relaxation characteristics), temperature, strain rate, etc. A typical stress-strain characteristic of M D P E described by Stewart et al (1999) is shown in Figure 3.3(a). As shown, the stress-strain behavior of M D P E can be categorized into three regions. This different behavior in each category depends on the molecular response in each stage of loading. In general, crystallinity of the molecular structure wi l l result in elastic-like behavior, while amorphous structure would cause polymers to behave as a viscous fluid. The stress-strain curve shows linear elastic behavior at small strains (instantaneous recoverable strains) followed by viscoelastic (recoverable strains but delayed recovery) 36 and viscoplastic (inelastic response, irrecoverable strains) characteristics. For many other plastics, it has been difficult to define a specific yield point or proportional limit for M D P E . In general, M D P E can stretch more than 400% of the tensile yield stress before breaking. Effect of strain rate and temperature on stress-strain behavior The stress-strain behavior of M D P E pipe is very complex due to temperature and strain-rate dependencies. Figure 3.3(b) and (c) shows the effect of strain rate and temperature on stress-strain behavior. A s discussed by Stewart et al. (1999), the modulus of P E significantly depends on the variation of temperature, e.g. between 0-49°C, the modulus can vary by a factor of two with higher modulus at lower temperatures. As shown in Figure 3.3(b), with increasing strain rates, the modulus wi l l increase, resulting stiffer stress-strain response. A t higher strain rates, the molecules have less time to orient to the direction of loading, thus results in a larger modulus. As reported by Suleiman and Coree (2004), the strain rate would have significant impact on stress-strain response for strain rates between 10"1 to 10"5 / s. 37 Zone I: Stretching of Inter-Atomic Bonds Zone H: Straightening of Molecular Chains • / \ A A A if) a> CO Zone JH: Relative Displacement of Molecules Strain, e (a) t o ^_ Incr. Temperature X Incr. Strain Rate t o CD o5 Figure 3.3. (a) Typical stress-strain behavior of Polyethylene and the effect of (b) temperature (c) strain rate, on the stress-strain behavior of Polyethylene (Stewart, et al, 1999). 38 Figure 3.4 shows the strain-strain curve derived from compression tests conducted on P E pipe specimens by Anderson (2004). The specimens were uniaxially compressed at a displacement rate of 1 mm/min, which was comparatively similar to the displacement rate of the pullout tests performed during the current series of tests (0.6 mm/min). 20,000 T i ; 1 1 ! 1 0.14 Strain Figure 3.4. Stress-strain characteristics derived from compression tests conducted on 60-mm and 114-mm P E pipes (Anderson, 2004). 39 3.1.2.3 The pipe diameter used in this study Four pipe sizes have been employed in different pipe configurations (see Figure 3.5). Table 3.1 presents the properties of these pipe sizes. The branch pipes (16 mm and 42 mm dia) and tees are attached to the trunk line (60 mm and 114 mm dia) using electro-fusion joining technique. The performance of the joint depends on the quality of the electro-fusion joint. In this test series, the pipes and tees were heat fusioned by experienced personnel of Terasen Gas, and it is fair to assume that they represent typical field tee connections. Table 3-1. Details of P E pipes used in testing Pipe/Tube diameter (mm) Wal l thickness (mm) M . O . P * (kPa) S .D .R* * 114(4") 10.32 690 (100 psi) 11 60 (2") 5.48 690 (100 psi) 11 42 (1 1/4") 4.22 760(100 psi) 10 16 (1/2" tubing) 2.28 n.a n.a * M.O.P: Max imum Operating Pressure * * S.D.R: Standard Dimension Ratio (Minimum outside diameter/Minimum wall thickness) 40 3.2 Details of the apparatus/ test setup The pipe-soil interaction testing facility at the University of Brit ish Columbia (Anderson et al, 2003) as shown in Figure 3.6 and 3.7, was used for the present study. The facility mainly comprises the fol lowing: 1. A test chamber (soil box); 2. Two actuators and a hydraulic control system; 3. Associated instrumentation and; 4. Data acquisition system 3.2.1 Soil box The test chamber has dimensions of 3.8m (L) x 2.5m (W) x 2.5m (H), and it permits conducting longitudinal and lateral pipe pullout tests (see Figure 3.5 for typical pipe test configurations.) A s noted by Anderson (2004), the dimensions of the soil box were designed considering the following: 1. The soil box should be long enough to permit sufficient relative soil movement along the length of the pipe during longitudinal pullout tests on straight pipes. Furthermore, in lateral pullout tests, the length of the box should be sufficient to develop both passive and active soil wedges (i.e. soil wedges in front and rear of pipe) without interfering with the walls of the box. 41 2. The width of the box should accommodate a reasonable length of branch pipe, to study its deformations and anchoring characteristics. In lateral pullout tests, the width of the box should be sufficient to achieve a reasonable amount of flexural deformation (mainly due to bending) for the considered pipe diameters. The walls of the box have been stiffened with 90 mm x 140 mm lumber cross beams (rib beams) spaced vertically at 300 mm intervals. These beams were supported by a steel frame, comprise of steel uprights (H-sections) which were bolted to the strong floor of the building. The steel uprights supports 19-mm ply wood panels which were covered with smooth stainless steel sheets from inside. These design details would minimize the outward deflection of the wal l , thus maintaining essentially no lateral movement of soil mass at the boundary of the box. 42 Branch Pipe Figure 3.5. Typical pipe configurations of (a) axial, (b) lateral and (c) branch pipe pullout test Smooth stainless steel sheets were used to line the side walls to reduce the friction between soil and the wal l . The interface friction angle for stainless steel and sand has been found to be around 20° (Brachman, et al, 2000, Karimian, 2006). In general, due to the large scale of the box, the wall boundary conditions would have minimal impact on the test results. Further details of the sand box are given in Anderson (2004) and assessment o f potential test errors are discussed by Karimian (2006) and it is not necessary to repeat them in this document. The frictional resistance of the gasket During axial pullout of straight and branch pipes, the pipes would pass in and out of the box through opening(s) on the walls o f the sand box. A gasket made out o f 50-mm thick hard rubber ( M L C foams) allowed the passing of the pipe through these openings without loss of sand from the box, while providing minimum friction against pullout. As mentioned by Anderson (2004), no lubricants were used, since that could increase the resistance at the interface and reduce sand particles sticking on to the rubber. Gaskets were not provided at the outlets of 16 mm branch pipes, given that the small access hole was considered to be adequate in prevent sand from escaping. However, it was ensured that the hole was slightly larger than 16 mm, so that additional friction is not generated at the outlet. 44 -380crt-String Potentiometers --eOOcrr Figure 3.6. Plan view of the soil box and the test setup flotentiome~t£,r Figure 3-7. Side view of the soil box and the test setup Separate pullout tests were performed to find the gasket friction. The test results showed that the maximum frictional force at the gasket would be around 0.05-0.1 k N in both 114 mm and 60 mm pipes. This value is marginally above the noise level of the load cell. Thus, it was judged reasonable to assume that gasket does not contribute to the pullout resistance of the pipe. 3.2.2. Actuators and hydraul ic system The pipe configurations were tested using a system of double-acting hydraulic actuators capable of applying a constant rate displacement at a selected end of the pipe configuration. These hydraulic actuators were manufactured by Royal Westcoast Cylinders Inc., N e w Westminster, B .C. , Canada. Each actuator has the capacity to impart 418 k N (93 kips) at 21 M P a (3000 psi) maximum working pressure of the in-house hydraulic system. The maximum displacement achieved from the actuator (stroke) is limited to about 600 mm. The actuators were trunnion-mounted to cast iron loading pedestals. The pedestal was bolted to the 0.6 m (2') thick structures lab reinforced strong-floor with 37 mm (1 1/2") steel bolts. The trunnion axis (same as the springline o f the pipe) was located at a level o f 0.8 m above. 47 The control systems of the actuators consist of: • Delta R M C controller, model R M C 1 0 0 - S 2 - E N E T , manufactured by Delta Computer Systems Inc., Vancouver, W A , U S A . • PQS servo Filters, with working pressures of 3000 psi and proportional valves manufactured by PQ systems Ltd, Burnaby, B .C. • SSI feedback probe, Temposonic R P with sliding magnet Using the R M C W i n software, it is possible to control the displacement, velocity and acceleration in four axes. The sliding magnetic SSI probe that is connected to the actuator cylinder provides the feedback to the Delta R M C controller. Thus, by adjusting the servo valve to control the l iquid flow, the target position of the actuator can be achieved. Furthermore, using the Delta R M C controller, the movement of the actuators could be synchronized, when two actuators are used in lateral pullout tests. However, this was not considered necessary as the operating speeds were very small. Instead, the starting commands were given to the two actuators separately. The differential movement arising due to disparity in starting time is insignificant. Couplings In axial pullout of straight and branch pipes, two piece split-clamps (made in-house) were used to connect the pipe to the actuator (see Figure 3.6.). In some tests, i f the anticipated strains were large, there was a possibility that the pipe may slip from the coupling due to 48 large diametric reduction of the pipe. In these situations, it was required to regularly tighten the nuts of the spilt-clamp to ensure no slippage. Figure 3.8 Two piece split- coupling used at the leading end of the 114 mm pipe. In lateral pullout tests, the pipe was connected to the actuator through two 3.6 m long high strength steel cables of diameter 15 mm. Because of the high stiffness, the extension of this cable is negligible under the applied levels of loading. A shackle connects the steel cable to a clamp at the pipe end. The other end of the steel cable was connected to a steel 49 lug which was fixed to the actuator head. This steel cable passes through a (120 mm x 50 mm) vertical slot cut in the front wall of the soil box to provide enough clearance between the steel cable and top of the slot, as the cable could move upwards in harmony with the pipe during lateral pullout tests. Hard rubber foam was used to cover this opening. 3.2.3 Instrumentation 3.2.3.1 Measurement of pullout resistance using load cells In axial pullout tests on straight and branch pipes, a 90 k N (20000 lb) capacity load cell was used at the leading end of the pipe (Load Cel l Type I). In lateral pullout tests, two load cells of 225 k N (50000 lb) capacity were employed in each pull ing end (Load Cel l Type II). A s a result of expected smaller loads in the branch pipes, load cells with a capacity of only 22 k N (5000 lb) were used in branch pipe tests, when pipes were restrained from movement (Load Cel l Type III). Details of these load cells are given below, • Load Cel l Type 1: Capacity 90 k N , (20000 lb), Type U - l . o Manufacturer: Baldwin-Lima-Hamilton Corp, Waltham, U S A • Load Cel l Type II: Capacity 225 k N , (50000 lb), Model 661.22 o Manufacturer: M T S systems corporation, Minneapolis, Minnesota, U S A 50 • Load Cel l Type III: Capacity 22 k N , (50001b), Model BTL -FF62 -CS-5k o Manufacturer: Transducers, Inc, Santa fe Spring , California, U S A 3.2.3.2 Measurements of pipe displacement using string potentiometers The displacements of selected locations of the pipe were measured using string potentiometers having a range of 2.5 m (100") and a resolution of ±2.1 mm. A 0.5 mm-diameter high strength steel string potentiometer wire was connected to a 1-mm dia 10-mm long screw which was driven into the P E pipe at selected locations. The resistance created by the screw and the string potentiometer wire was low and assumed to be negligible. 3.2.3.3 Measurement of pipe strain using strain gauges As shown in Table 3.4, most of the pullout tests performed were strain gauged. The selection of the type of strain gauges (see Table 3.2 for details of strain gauges) and procedure for mounting strain gauges has been reported in detail by Anderson (2004). A similar procedure with some modifications was followed in this study. In all tests, an array of strain gauges was attached to selected locations along the crown of the pipe. In some tests, additional gauges were attached at the pipe invert level to obtain the strains at two locations of a selected cross section. The excitation voltage for string potentiometers and load cells was 10 V , and 5 V for strain gauges. 51 Table 3.2. Details of the strain gauges Manufacturer Kyowa Type KFEL-5 -120-C1 Gauge Length 5 mm Strain Range ± 15% Gauge Resistance 119.8 ±0 .2 . Gauge Factor (24°C, 50%RH) 2.13 ± 1.0% Temperature Coefficient of Gauge Factor + 0.015 %/ °C 3.2.4 Data acquisit ion system The test data pertaining to pullout resistance (load cell), displacement (string potentiometer) and strain in the pipe (strain gauges) were collected using an in-house made signal conditioner and a digitizer. The channels in the signal conditioner was limited to 16, hence, in some tests, the number of strain gauges and string potentiometers were limited accordingly. During experiments, the resistance, displacement and strains were monitored in real time using GUI interface provided by the commercially available software, D A S Y L A B (Data Acquisit ion System Library, 2004). The data was recorded at a rate of 1 samples per second to match the slow displacement rate, thus to limit the size of the output file to a manageable size. However, data was further preprocessed by filtering to 1/10 t h o f its size, before performing any data manipulations. 52 3.3 Preparation of test specimens This section presents the following key components related to preparation of specimen for testing (a) Mounting of strain gauges (b) Placement of pipe in sand box (c) Placement of sand 3.3.1 Mounting of strain gauges on PE pipe In this process, the selected location of the P E pipe was first marked and cleaned with a soft cloth. Afterwards, the surrounding area (encompassing approximately 3 " diameter area) was etched with a 220 grit sand paper. Since P E has a soft surface texture, excessive applications of sand paper would cause deep indentations on pipe surface. Since this excessive roughness could lead to poor bondage between strain gauge and pipe. Care was taken to ensure that the pipe surface is only slightly roughened. Next, the dirt and other contaminants were cleaned (wiping in one direction) using a degreaser and the surface was allowed to dry for about 15 min. The dried surface was then applied with a coat of Loctite 770 primer and allowed to dry for about half an hour. The strain gauges were then positioned and aligned in the required direction, with the aid of a cellophane tape. If required, the tape was adjusted as required, so that the gauge is inline with the intended 53 direction. For relatively smooth M D P E surfaces, Anderson (2004) demonstrated that strain gauge cement such as "Loctite 414 instant adhesive" is sufficient to provide a good bond between strain gauge and the pipe. This strain gauge cement was applied sparingly, underneath the gauge by pealing the cellophane tape from one end. Pressure was applied on the gauge using the thumb for 1-min to ensure that excess bonding agent was expelled. The gauges were then secured by placing a rubber and a steel pad on top of the gauge and strapping them to the pipe with an electrical tape to maintain a nominal pressure while setting. After elapse of 24 hours, the pads were removed and a thin coat of "M-Coat A " , air drying polyurethane coating was applied on top of the strain gauges and allowed to dry for few minutes. The terminals of the gauges were then soldered to the soldering pad which was also secured to the P E pipe using the same adhesive. After soldering the wires, the resistance of the gauges was measured to ensure proper circuit. A stress relief loop was formed from the wires and secured to the pipe as shown in Figure 3.9. The solder pad and the wires were further secured to the pipe with a rapid setting epoxy. Excessive amounts of epoxy were avoided as the epoxy could locally strengthen the pipe. Alternatively, epoxy would help to prevent sand abrasion and consequent damage to the solder pad. After the epoxy has dried, nylon cable ties (290 mm long) were used to secure the wires in place. In dense sand test, 4-5 cable ties were needed, depending on the location of the strain gauge. The gauges that were placed immediately after the tee, requires less strapping. Besides, it was found that four cable ties were enough in loose tests to obtain strain gauge readings up to a reasonable extent of 54 pipe displacements. Final ly, the gauge area was further strapped with an electrical tape. The entire strain gauge area may extend up to a length of 110 mm. Figure 3.9. Formation of stress relief loop from lead wires of the strain gauge. Except for few gauges in the 60-mm branch pipe tests, none of the gauges detached from the gauge-pipe interface during testing, despite significant abrasion at the pipe-soil interface. This observation substantiated that proper bonding was achieved between pipe and the strain gauge. In the tests with 60-mm dia. pipe, the pipe underwent very large strains (maximum measured was 7%); in such situations, it was noted that the shear strength capacity of the adhesive was exceeded. Instead, in initial tests, it was observed 55 that the signal from the gauges seized due to the damage to the soldering pad. Therefore, in later tests, excess amounts of solder and larger solder pads were used to strengthen the connection between the strain gauge terminals and the wires. However, even with these additional protections at the soldering pad, the reliable strain gauge readings can only obtain up to a displacement 100 mm of the leading end, in dense sand. After completion of the above strain gauge installation, at selected locations, the pipe was ready for placement inside the soil box. The wiring was done in such a manner to prevent possible increase in noise level and to minimize the drag effects caused by the wires. The occasional spikes showed in measurements were presumed to be caused due to dragging of these wires. 3.3.2 Placement of pipe in sand box The pipe configurations (axial, lateral and branch pipe) were placed in the sand box, as shown in Figure 3.5(a) to (c). In axial and branch pipe pullout tests, due to large scale of the box, the impact on test results from wall boundary conditions wi l l be minimal. In lateral pipe pullout tests, sufficient clearance was given in front and back of the pipe to allow development of passive and active soil wedges without interfering with the front and back end of the box (see Appendix B for pipe configurations). A t the start of the test, the pipe would have a clearance of 2.4 m in front and 1.4 m to the rear wall. Karimian (2006) showed through numerical modeling and simple friction calculations that the dimensions of the soil box were adequate to perform lateral pullout tests without any 56 interference from the walls, even for larger diameter (457 mm) steel pipes. Furthermore, as mentioned in Section 4.2 due to bending of the pipe, the surface deformation closer to the sidewalls was small. Therefore, the sidewall friction could be assumed to be insignificant compared to the overall lateral pullout resistance. A lso , the pipes were placed inside the soil box, such that there was sufficient clearance (-25 mm) from the sidewalls. This could avoid pipe ends coming into contact with the sidewall as the pipe bends (due to possible increase in pipe length). 3.3.3 Placement of sand Sand was stored in bulk bags (with discharge chutes), in which each can carry up to 1 m 3 of sand. About 19 m of soil was needed for each test. Since all tests had a burial depth of 0.6 m, the volume of sand needed for each test was approximately the same. The sand bags were moved to the inside of the sand box using an overhead crane. B y untying the knot and opening the chute at the bottom, the sand was spread uniformly as possible while maneuvering the bags with the crane. The drop height was roughly maintained at 75 mm to 100 mm. For the dense sand tests, sand was placed in ~ 125 mm lifts and each lift was compacted using '/2-ton roller with four passes each in lateral and longitudinal directions. In loose sand tests, the leveling of sand was avoided to ensure minimum possible level of disturbances to the sand and its structure. Leveling was done only when sand was filled 57 up to the required level. Tests were performed within 24 hrs after f i l l ing the box so that any effects due to ageing of sand would be similar in all tests. 3.3.3.1 Density of soil backf i l l Anderson (2004) and Karimian (2006) have already developed soil compaction procedures to obtain specimens having reasonably uniform density. This involved placement of sand in about 125-mm lifts and compacting each lift to desired density. Anderson (2004) used a nuclear densometer to measure the density, while Karimian (2006) also used a nuclear densometer in some tests and density pan measurements in other tests. Figure 3.10 shows the density measurements obtained by Anderson, using a nuclear densometer on dense and loose sand. These measurements revealed a relatively large scatter in density in loose sand and the scatter become small when sand is compacted. These same compaction procedures were adopted for the present study. In all tests, the density was obtained by pushing a sharp cylindrical sampler (diameter 75.9 mm, height 70.3 mm and wall thickness 1.5 mm) into the sand. The density was calculated by dividing the weight of the sand extracted from the sampler by the volume occupied. In each test, five measurements were taken from the surface, at different locations. The calculated density measurements showed an average density of 1610 kN/m and a standard deviation of 0.15 among dense sand tests and an average of 1531 k N / m 3 and a standard deviation of 0.09 for loose sand tests. The average densities for each of these 58 tests are discussed in Table 3-3. These density values were consistent with the densities reported by Anderson (2004) and Karimian (2006) for the same sand. Besides, it is worth noting that in these tests, the same soil compaction procedure was adopted as in the case of Anderson (2004) and Karimian (2006). In addition, the insitu density was also measured using sand cone method, as stipulated in A S T M D1556-90 (1996). These test results were also consistent with the density pan measurements obtained after the test. Approximately similar average densities among tests indicate good repeatability, especially in dense tests. The soil density measurements obtained in these tests are considered accurate and acceptable, in view of the variability of soil density that is l ikely to occur in these types of large scale tests. 59 10 8 + > 6 -c V 3 u_ 4 --2 f U I 1 i 1 i 1 i ' i ' i ' i ' i i ' i i i i i i i i i i i i i i I i i i i r i i i i i ' ' I 1350 1400 1450 1500 1550 1600 1650 1700 Density (kg/m3) (a) 10 o u c V 3 O -V £ 4 2 4 1350 1400 1450 1500 1550 1600 1650 1700 Density (kg/m3) (b) Figure 3.10. Density measurements using nuclear densometer in (a) dense and (b) loose sand tests (Anderson, 2004) 60 Table 3-3. Average densities measured in tests. Test number (dense) Density/ (kgm3) 60D 16.32 60D-1 16.34 114D 16.01 114D-1 16.00 114D42-U 16.19 114D42-R 15.93 114D16-R 16.17 114D16-U 15.96 60D16-U 16.13 60D16-R 16.17 114D-(1.5m) 15.87 60D-(tee only) 16.06 Average density (dense) 16.10 Test number (loose) Density/(kgm"3) 114L42-R 15.37 114L42-U 15.24 Average density (loose) 15.30 After these test preparations mentioned above, the test was started with pulling of the pipe from one end (referred to as the leading end) by the actuator while recording the 61 pullout resistance, strain and the displacement at selected locations. For the selected rate of displacement (36 mm/hr), a single test may run up to more than 16 hours to achieve a displacement of ~ 600 mm, at the leading end of the pipe. Emptying of sand box and dust control aspects A 3-m tall screen was erected at top of the box to control dust during specimen preparation. The screen can be lowered by a string-pulley arrangement, when required. The tarpaulin screen was supported by steel struts of 55 mm (W) x 3 mm (T) hollow square sections, at the four corners of the box. Once the test is completed, the sand inside the box was emptied through an access port of 450 mm x 300 mm, at the bottom of the side wall. The sand flows through an exit chute made out of a geomembrane, to an inclined conveyor belt (10 ft long). A sand bag was placed at the other end of the conveyor belt to collect the sand that is coming through the conveyor (see Figure 3.11). To control dust, a moveable chamber was placed, covering the conveyor and other arrangements. This chamber was made out of 1 l /2 "x l 1/2" wooden frame and covered with transparent polythene sheets. A n exhaust fan was placed, closer to the sand bag (where the dust generation is at its most) to remove the dust from the chamber. A plastic duct leading from the 14" exhaust fan blows away the dust that is accumulated inside the chamber. 62 Figure 3.11. Arrangement for unloading of sand from the sand box 3.4 Test program A series of pullout tests were undertaken on select pipe configurations and soil density conditions. A test matrix, as shown in Table 3-4 (a) to (c) was developed. This included eight branch pipe tests, four longitudinal pullout tests, and two lateral pullout tests. As may be noted, in addition to above tests, several supplementary tests were performed to examine different interaction parameters. These include two tests performed after removing the branch pipes (having tapping tees only), a branch pipe test conducted to evaluate the performance of the butt-tee connection and longitudinal pullout test on a 63 concrete fi l led pipe. A lso two trial tests were performed, initially, but the results wi l l not be presented in this thesis due to uncertainties in the load cell measurements. In order to facilitate data review and discussion, the tests were identified with a numbering system as described below. The test numbers were attributed so that they would readily reflect the trunk pipe diameter, branch pipe diameter, soil density, and the end boundary condition of the branch pipe. The number in front of the first letter denotes the diameter of the trunk pipe and the number following this letter stands for the branch pipe diameter. The characters " D " and " L " refers to cases having dense and loose sand backfill conditions, respectively. " U " refers to the tests where branch pipes were kept unrestrained whereas, " R " designates branch pipes were restrained against movement, e.g. 60D16-U refers to 60 mm diameter trunk pipe with unrestrained 16 mm branch pipes, in dense sand. A lso note that the wildcard character " * " w i l l be used here to denote a test series, e.g. 60D16-* wi l l refer to 60D16-U and 60D16-R tests. Table 3-4. Test matrix (a) Longitudinal (axial) pullout tests Test ID Pipe diameter (mm) Strain gauged or not Soil density 60D 60 No Dense 114D 114 No Dense 60D-1 60 Yes Dense 114D-1 114 Yes Dense 64 (b) Lateral pullout tests Test ID Pipe diameter (mm) Length of the pipe (m) Soi l density 114D-(2.5m) 114 2.5 Dense 114D-(1.5m) 114 1.5 Dense (c) Branch pipe pullout tests Test ID Trunk pipe diameter (mm) Branch pipe diameter (mm) End condition of the branch pipe Soi l density 60D16-U 60 16 Unrestrained Dense 60D16-R 60 16 Restrained Dense 114D15-U 114 16 Unrestrained Dense 114D15-R 114 16 Restrained Dense 114D42-U 114 42 Unrestrained Dense 114D42-R 114 42 Restrained Dense 114L42-U 114 42 Unrestrained Loose 114L42-R 114 42 Restrained Loose 65 (d) Supplementary tests Test ID Description 114D42(tee only) Pullout test to find the resistance of the tapping tee 60D16 (tee only) Pullout test to find the effect of the tapping tee 114D42-U (butt tee) Branch pipe test on 114 mm trunk pipe and 42 mm branch pipe, connected via a butt tee (unrestrained) 114D(concrete) Ax ia l pullout test on 114 mm pipe fi l led with concrete AD60-1 Trial test on 60 mm pipe without strain gauges AD60-2 Trail test on 60 mm pipe with strain gauges 3.5 Selection of test parameters 3.5.1 End boundary conditions for the branch pipe In these full-scale tests with P E pipes, selecting a proper end boundary condition for the branch pipe is imperative, since the restraint condition could affect the overall strain distribution in the pipe system. Furthermore, in certain pipe configurations, the failure mode could also be affected by the end boundary condition of the branch pipe. Generally, in field installations, the small diameter branch pipes wi l l be laid with some slack (snaking), allowing some movement of the pipe. On the other hand, the branch pipes can 66 also be significantly restrained by pipe components (joints, bends etc) or by the longitudinal frictional anchorage or by any physical restraint exist in house connections, as shown in Figure 3.12. Thus, to simulate these field conditions, the full-scale tests were conducted with and without end restraint conditions. In large diameter branch pipes, the branch pipe is more likely to be restrained, than not, under general field conditions. Steel pipe (riser) —* Ground surface Plastic saddle fusion tapping tee Protective sleeve <^x<V X P V C protective sleeve Plastic to steel transition coupling Plastic main Figure 3.12. Typical field installation of a distribution pipe system (details extracted from office of pipeline safety, 2005, http://ops.dot. gov/). 3.5.2 Burial depth To prevent exposure to U V radiation and construction damages, most distribution pipes are installed with a soil cover ranging between 0.3-1.5 m. With the knowledge of this and 67 to limit depth variable, in all tests, the burial depth was selected to be 0.6 m. Thus, the H/D ratios of this study correspond to 5.2 and 10 for pipe diameters of 114 mm and 60 mm, respectively. (Note that, H is the burial depth and D is the outer diameter of the pipe as defined in Section 2.2.1). 3.5.3 Rate of ground displacement For full-scale testing, a suitable rate of displacement needed to be selected as the overall stress-strain response would depend on the applied strain rate, as mentioned in Section 3.1.2.2. According to detailed landslide monitoring in Italian mountainous areas, Bughi et al, (1996) reported that 90% of the ground movements can be categorized as "extremely slow" (according to Varnes Classification, 1958) and majority of them are confined to a depth of 6 m below the ground level. These average ground movements can be as slow as 10" - 10" m/yr (0.16 mm/day), with seasonal fluctuations due to precipitation and snow meltdown conditions, etc. These seasonal effects could induce rapid displacements, in turn inducing potentially large pipeline stresses in the short term. In the current series of tests, the applied rate of displacement of the leading end of the pipe was selected to be 36 mm/hr. The selected rate was the slowest rate of displacement that can be practically achieved from the hydraulic controller system. Although the selected rate is faster than the average slope movements in landslides, the rate is considered reasonable, taking into account the practicalities and the relatively fast rates of ground movements that are possible during seasonal fluctuations. Except in some tests 68 where the displacement rate was increased after reaching a certain displacement (see Section 5.6), the main focus was to investigate the pipe pullout response under the selected constant displacement rate of 36 mm/hr. 3.5.4 The testing temperature As pointed out earlier, the modulus of the pipe depends on the operating temperature. Moreover, the creep behavior of the P E wi l l also influenced by the field temperature, e.g. lower temperature wi l l give rise to higher stiffness modulus and lower creep deformations. Therefore, all tests were performed at an ambient laboratory temperature of 20°C, thus any temperature effects among tests were insignificant. 69 CHAPTER FOUR A X I A L A N D LATERAL PULLOUT TESTS ON STRAIGHT PIPES Understanding the basic soil interaction in pipe components (axial movement of the trunk pipe and lateral movement of the branch pipe) is imperative in analyzing more complex pipe configurations. Wi th this in mind, several axial and lateral pullout tests were performed on straight lengths of pipes. In contrast to steel pipes, the interaction behavior in P E pipes would depend on many factors. The higher flexibil ity and viscoelastic stress-strain behavior of P E pipes would result in different force-displacement (soil-spring) characteristics that are different in comparison to those for steel pipes. This chapter discusses the experimental observations and the resulting analysis of the axial and lateral pullout tests on buried straight lengths of P E pipes. For axial loading situations, a closed-form solution was developed on a preliminary basis considering the nonlinear material behavior of P E , and the derivations related to this solution are presented. The test results of the axial pullout tests were also compared with other pipe-soil interaction models published in literature. Test results are also reviewed with regard to the impact of cross-sectional area changes in the pipe due to Poisson's effect. Several factors that would limit understanding the pipe-soil interaction, especially 70 when the soil movement is along the pipe axis and these limitations are discussed in detail, as a part of this chapter. A lso in this chapter, the results of two lateral pullout tests wi l l be presented, and the results wi l l be used in potential future numerical modeling of lateral pullout behavior o f the p ipe. . 4.1 A x i a l pul lout tests Four axial pullout tests were performed on straight pipes having diameter of 60 mm and 114 mm, and the corresponding details are given in Table 3-4. In all tests, the soil was compacted and the resulting average soil density was 1610 kg/m (75 % relative density). In addition, to 60D and 114D tests performed on bare pipes, another two tests (60D-1 and 114D-1) were conducted with strain gauges attached to the buried length of the pipe. The test configurations are schematically given in Figure 4.1. The results for individual tests are presented in Appendix B. 71 S P 2 <——| SP1 •4 SG1 L" . j I I J S G 2 Load Cel l / P E pipes with or without strain gauges. Figure 4.1 Typical pipe configuration for axial pullout behavior of 60 mm and 114 mm bare pipes. N O T E : The abbreviations " S G " and " S P " stands for strain gauges and string potentiometers respectively. 4.1.1 Observed load-displacement-strain response for axial loading Understanding of the load-displacement characteristics is an important consideration in the development of a soil-spring model for the axial movement of P E pipe. The variation of pullout resistance with respect to the displacement of the leading end of the pipe (measured using string potentiometers) for the 60-mm and 114-mm pipes is shown in Figure 4.2 (i.e. Test No . 60D and 114D). As explained in Section 3.2.1 the pullout resistance was not corrected for gasket friction, since this component was considered negligible. 72 In these tests, the displacements of the leading and the trailing end of the pipe were measured using string potentiometers (SP) attached to these locations of the pipe, and the corresponding results for 60-mm and 114-mm pipes are shown in Figure 4.3. The strain distributions along the pipe during axial pullout testing of 114-mm and 60-mm pipes are shown in Figure 4.4. The strain distribution is observed to be nonlinear along the length of the pipe, until the peak resistance is mobil ized. Thereafter, the strain distribution would become nearly linear. Similar observations have also been made by Anderson (2004). In addition to the axial strain along the buried length of the pipeline, the axial and the radial (hoop) strains at the leading end of the pipe (see SGI and SG2 in Figure 4.1) were also measured and those results are presented in the Appendix B. These radial and axial strain measurements were further verified through regular, independent caliper readings obtained during testing. 73 8 1 H 0 4 — , — i — i — i — i — i — i — i — i — | — i — i — i — i — | — i — i — i — i — i — i — i — i — i — | — i — i — i — i — ; 0 100 200 300 400 500 600 SP1 Disp lacement , x (mm) Figure 4 .2 Load-displacement characteristics for the 114 mm and 6 0 mm pipes, (refer to Figure 4.1 for the location of SP1). .40 J SP1 D i s p l a c e m e n t (mm) Figure 4.3 Displacement of the leading and trailing ends of the 114 mm and 6 0 mm pipes, (refer to Figure 4.1 for the location of SP1). 74 Figure 4.4 Strain distribution along the buried length of the 60-mm diameter pipe at different displacements of the leading end, (see Figure B.2.1 for strain gauge locations). To verify the validity of the strain gauge measurements, in some locations of the pipe, two strain gauges were attached to the crown and the invert of the pipe. A s observed in Figure 4.5 identical strains were observed, indicating that for a displacement of about 100-mm of the leading end, the strain gauge would yield very accurate readings. As a general observation, although the sizes of the test box was different, as a whole, the pullout test results of branch pipes and straight pipes were in agreement with the results obtained by Anderson (2004). 75 12000 0 20 40 60 80 100 120 140 SP1 Displacement (mm) Figure 4.5 Strain gauge readings for 60 mm straight pipe at a distance of 0.75 m from the leading end. The axial tests conducted on straight P E pipes indicated that the pipe displacement (x m a x ) required for mobil izing the peak soil resistance is in the range of 20-30mm. This is significantly larger than the x m a x values of 2 to 3 mm observed from tests conducted on steel pipes ( A S C E , 1984, Karimian, 2006). In steel pipes, owing to larger stiffness and limited pipe length inside the test chamber, pipes would essentially act as a rigid element; hence, essentially no contribution from the elastic/plastic elongation of the pipe to x m a x . In P E pipes, due to relatively smaller interface friction angle (friction coefficient of 0.6) compared 0.8-0.9 of steel pipes, the full friction at the interface would only be mobilized at a larger relative displacement than in steel pipes. 76 In P E pipes, the frictional resistance developed at the interface would cause the pipe to elongate, as may be noted in the strain gauge readings in Figure 4.4. The mobilized frictional length increases progressively with increasing leading end displacement of the pipe. As a result, x m a x would depend on pipe properties (pipe cross sectional area, pipe length tested), material properties (stress-strain behavior) and soil characteristics (burial depth, density, friction angle, lateral earth pressure coefficient) as wel l . In other words, defining a unique x m a x value as in steel pipe is not appropriate for P E pipes, unless the pipe's elongation is small to assume that the pipe would behave as a rigid element for the considered length. Therefore, above P E pipe pullout tests should be considered as "model tests" not "element tests" as would be the case for tests on rigid steel pipes, where the force per unit length is calculated by dividing the total axial force by the total length of buried pipe. 4.1.2 Analysis of axial pullout resistance - Closed form solution The axial pullout testing conducted on straight pipes provided an opportunity to investigate the nonlinear material behavior o f P E pipes, and to develop a closed-form solution to account this stress-strain response. When the pipe configuration is subjected to differential ground movement, gradual softening of the deformation modulus of the pipe with increasing strain would generally result in smaller pullout resistance than a pipe with constant modulus. Therefore, i f the nonlinear stress-strain behavior is not taken into account, the pullout resistance wi l l be overpredicted in P E pipes subjected to ground displacements along the pipe axis. The derivation is based on the "theory of beams on 77 elastic foundation" (Hetenyi, 1946). In this formulation, for axial loading situations, a second order differential equation can be derived and solved for known boundary conditions. The derivation is based on the force equilibrium at the pipe-soil interface. A set o f equations were developed for a slide geometry in which the soil mass surrounding the buried pipe is stable (further aspects regarding the slide geometry is discussed in Section 2.3). A s per the analysis, the pipe subjected to differential ground movement is considered, hence it could be located in the moving or in the non-moving soil block in which the soil in the immediate vicinity of the pipe is stable. It is judged that this approach would simulate typical failure mechanisms likely to occur under slow ground movements, wherein the slip surfaces are located at a greater depth. 4.1.2.1 Pipe-soil interaction for an element level pipe length At the interface, the pipe-soil interaction is assumed to fol low an elastic-perfectly plastic response as shown in Figure 4.6. From direct shear tests performed on PE-sand interface (O'Rourke et al , 1990), the maximum elastic displacement at the interface (u x) was considered to be around 0.5 mm in dense sand. A s mentioned in Section 2.1.1, the simplified Equation 2.1 is commonly used in determining the axial soil-spring characteristics for steel pipes. A lso in this interaction formulation, the maximum resistance per unit length (T m a x ) of pipe could be obtained from the Equation 2-1. As discussed in Chapter 2, the selection of appropriate parameters for K and cp is a difficult task particularly due to the combined variability of friction angle of soil, normal soil 78 stress on the pipe and the diametric reduction of the P E pipe. Despite these difficulties, for the illustration of this model, K and cp were taken to be 1.6 and 39° for 60-mm pipe, respectively. The rationale for selecting these values for modeling wi l l be further discussed in Section 4.1.3. 4.1.2.2 Extension of element level interaction to model pipe response in a physical soil box test The above model representing the pipe-soil interaction at the element level was applied to simulate the response of a straight P E pipe during axial pullout loading tests described in Section 4.1. Depending on the level of relative axial displacement, separate sets of equations can be derived considering the following two regions of interaction: (a) Region 1: Initial pipe and interaction response (u < u x) and, (b) Region 2: When interaction response is plastic and pipe material response nonlinear (u > u x) with respect to Figure 4.6. Note that in the following formulation, the "relative" displacement between pipe and the soil wi l l be denoted by " u " . Considering a typical pipe pullout test, the length of pipe that would have elements interacting in Region 1 and 2 is illustrated in Figure 4.7. 79 T m a x obtained using A S C E (1984) Equation 2.1 0) o S-l .a c » o >-<u 3 Region 2 Region 1 Relative displacement, u Figure 4.6 Interaction behavior at the MDPE-sand interface x = 0 Leading end of the pipe Pull ing force, Po > Figure 4.7 Mobi l ized elastic (x e) and plastic (x p) frictional lengths along the pipe. (a) Region 1 : Initial pipe and interaction response (u < ux) As the relative displacement (u) is less than u x (= 0.5 mm), the second order differential equation given in Equation 4-1 can be derived and the corresponding solution is given in 80 Equation 4-2, in which ui is the relative displacement for the Region 1, whilst A i and A2 are constants. d_ dx E,AP KdX ; [4-1] = A\erx + A2e -yx [4-2] I TCD T where, y = — and D is the pipe diameter and A p is the cross sectional area of the v E < A n u * pipe. This formulation is similar to that proposed by O'Rourke and Nordberg (1992), Trigg and Rizkal la (1994) and Rajani et al (1995), except that the use of subgrade modulus is avoided in place of ( T m a x / u x) as the initial gradient of the interaction response (i.e. frictional behavior at the interface). It is important to note that initial gradient of T vs. u in Region 1 would not necessary correspond to the subgrade modulus of the soil. In addition, for small relative displacements (u < u x), as the induced force in the pipe is small, the modulus of the pipe was assumed to be constant (i.e. the initial Young's modulus of P E , E i ) . This assumption is considered reasonable for pipe diameters (i.e. 60-mm and above) assessed in this formulation. 81 (b) Region 2: When interaction response is plastic and pipe response is nonlinear (u > ux) A differential equation similar to the previous case was derived for the elements of pipe that would fall into Region 2. A t this state, it is reasonable to assume that the maximum interface frictional force (T m a x ) would have mobilized in the pipe length x p (mobilized plastic frictional length) as shown in Figure 4.7. With regard to pipe material response of Region 2, the stress-strain behavior (which is initially assumed to be linear when u<ux) would show the typical nonlinear material behavior exhibited in viscoelastic materials, with increasing strain in the pipe. To model this stress-strain response, data from uniaxial compression tests performed by Anderson (2004) was available. A s shown in Figure 4.8, it was possible to approximate this response using a hyperbolic stress-strain relationship (Equation 4-3). In addition to these results, data from load cell and strain gauge readings mounted at the leading end of the pipe was available from axial pullout test conducted as a part of this study. These results are also presented in Figure 4.8 for comparison. There is a good agreement between the stress-strain curves derived for the P E material based on axial compression and axial extension modes. (Note that these uniaxial compression and pullout tests were performed at similar temperatures and displacement rates.) From above test results, the hyperbolic model constants " a " and " b " were selected to be 2.1xl0" 6.and 82 4.0x10"5 respectively for the related testing conditions (i.e. temperature and strain rate) with modeling units of kPa. 20,000 Strain Figure 4.8 Stress-strain response for the 60-mm M D P E pipe from uniaxial compression, axial pullout tests along with response modeled using hyperbolic model. [4-3] a + be Combining the Equation 4.1 to 4.3, the differential equation given in Equation 4.4 can be derived. The corresponding solution (for ui -the relative displacement in region 2) for 83 this differential equation can be obtained as in Equation 4.5, in which C i and C 2 are constants which depend on the pipe and interaction properties selected for modeling. a u Kdx2 y K a A P J a + b a — b du \dxjj [4-4] u2 = ' z + 1^  — - 2 z + log \ + \og(x + C[) + Ax + 2coC2 2a>\ \z-\J [4-5] where, (a) z = (-4Ax+ 1-4X0^ L2 rp (b) co = ^ [4-6] [4-7] (c) X = bAn [4-8] For ease of formulation, the origin of the axis was selected at elastic-plastic interface as shown in Figure 4.7. The remaining unknowns (C i , A i A 2 and x p ) of above equations (i.e. Equations 4.2 and 4.5) could be obtained by considering the fol lowing boundary conditions. (i) ) as x -> - 0 0 , u -> 0, in turn A 2 0; (ii) At x =0, u e = Up (the displacement is continuous); (iii) A t x =0, u e ' = Up' (strain is continuous); 84 (iv) At x = x p , P = Po (force at the leading end of the pipe); Although the boundary condition (i) would suggest an infinitely long frictional length for region 1, it was observed from strain distribution given in Figure 4.4, this mobilized frictional length (x e) would be around 0.5 m when the leading end displacement becomes 0.5-mm (the maximum displacement in region 1). This x e value of 0.5 m is proposed in determining the total mobil ized frictional length. Nevertheless, the impact of the above boundary condition has a smaller effect on the overall response of the pipe system since the impact (induced strain and force in the pipe) from the Region 1 is small for the pipe diameters considered in this study. With these boundary conditions, value of C i was obtained as 0.401 for the selected interaction and P E material properties. By differentiating the Equation 4.5, the strain along the pipe is derived as in Equation 4.9. If the formulation was considered in the context of full-scale test conducted herein, the corresponding force in the pipe at the leading end of the pipe (at a distance x p ) can be obtained from Equation 4.10. Moreover, from axial pullout tests performed on P E pipes, the resulting force-strain characteristics could be attained from the load cell readings and the strain gauges located at the leading end of the pipe. The experimental results and the predicted values from closed-form solution are plotted in Figure 4.9 for 60-mm pipe. 1 (2AX + Z + 2AQ - 1 ) [4-9] m(x + Cx) 85 P(xp) = E(xp)Apu'p(xp) [4-10] 5.00 60D 3 °- 1.00 co o CD £ 3.00 4.00 0.00 )P ' ' ' I I ' ' ' ' i 1 1 ' 1 0 2000 4000 6000 8000 10000 12000 14000 Axial strain (e) Figure 4.9 Predicted and measured (from Test No. 60D) force-strain relationship at the leading end of the pipe in 60-mm pipe. Another important outcome of this formulation is the ability to obtain a relationship between the mobil ized frictional length and strain (see Figure 4.10). Such relationship would provide a means of judging the performance of a pipeline subject to relative ground movement. 86 60000 Xp (m) Figure 4.10 Strain - frictional mobilized length (x p) relationship derived from closed-form solution. As can be observed in Figure 4.10, at a mobilized frictional length (x p) of 5.81 m, the strain in the pipe would become infinite and failure would occur due to yielding. The same value of x p could be derived from Equation 4.11, which was derived from Equation 4.6. Note that x p is the mobil ized frictional length in the region 2, hence the total mobilized frictional length would become x e + x p (= 0.5 + 5.81 = 6.31 m). x , = £ - C , [4-1U 87 Furthermore, by fitting a polynomial to Figure 4.10 and then integrating this polynomial would enable to obtain the displacement relationship for the region 2. It is important to note that Equation 4.5 could not be used directly to calculate the displacement due to numerical difficulties in solving this equation. Thus, from the polynomial approximation the relative ground displacement corresponding to x p of 5.81 (failure of the pipe) was obtained to be ~70 mm. This information is vital in assessing the performance of a buried pipeline subjected to ground movement. Using the data on pipe displacement, the force-displacement characteristics of the 60-mm pipe could be also derived, and results are shown in Figure 4.11. Note that for comparison purposes with the experimental results, this force-displacement relationship is plotted only up to a pipe displacement at which the trailing end of the pipe begins to move in Test No . 60D. Although, the pullout resistances were somewhat underpredicted in Figures 4.9 and 4.11 (the reasons are discussed in Section 4.1.3.), a fairly accurate representation was obtained from the closed-form solution. The slow rate of displacement may also have contributed to the underprediction, by inducing viscoelastic stress relaxations/ creep even during the test, thus relieving some stresses in the pipe. 88 5.00 SP1 Displacement, x (mm) Figure 4.11 Predicted and experimental force-displacement response in 60-mm pipe. 4.1 .3 Consideration in the selection of parameters to model for M D P E - s a n d interface at element level Many factors affect the pullout characteristic of M D P E pipe. Although the nonlinear material response of M D P E is taken to account in above formulation, the accuracy of such results is limited by the interaction parameters selected for the derivation of the above closed-form solution. As indicated in Section 4.1.2.1, the lack of knowledge in relation to variation of K (lateral earth pressure coefficient) and 8 (interface friction angle) during shearing is one of the 89 areas that would present difficulties in estimating the frictional resistance at the element level. As Karimian (2006) has noted, the estimation of K wi l l have greater impact on the pullout resistance compared to the interface friction angle. The formulation of the closed-form solution in Section 4.12 was undertaken using assumed pairs of K and 5 values. Therefore it is considered prudent to examine the range of variation of these parameters and comment on the basis for the selected K and 5 for the closed-form solution. (i) Selection of the interface friction angle (5) As given in Equation 2.2, it is common practice to denote the interface friction angle (8) as a function of the friction angle of sand (cp). From direct shear tests performed on M D P E pipes, it has been observed that the peak resistance at the sand-MDPE interface would be mobil ized after a displacement of 2-3 mm (O'Rourke, 1990). As discussed in Section 3.1.1.2, based on dry triaxial tests, the peak friction angle for Fraser River Sand was reported as 46.5° (Karimian , 2006), at an overburden stress of 10 kPa (the estimated overburden pressure at the springline of the pipe). Nevertheless, the mobilized friction angle at any instant would depend on the level of shear strain in soil. With increasing strain, due to soil dilation particularly in dense sand, the friction angle would reach a peak and then reduces to a residual value with increasing shearing. As observed in Figure 4.2, the larger x m a x (~ 20 mm) in axial pipe pullout tests would imply 90 that, at this displacement, the mobilized friction angle along some parts of the pipe should be between peak and the large strain friction angle of 33°. On this basis, although, a single value for cp is not likely to represent the entire length of mobilized shear resistance, the selection of an average friction angle of 39° is considered reasonable to compute the maximum pullout resistance in Figure 4.2 for the calculations in closed-form solution. (Consequently, the selected interface friction angle would become 23.4°, i.e. 0.6 times the internal friction angle of sand). (ii) Selection of lateral earth pressure coefficient (K) Despite the definition (ratio between horizontal to vertical soil pressure), in the context of using in Equation 2.1, the lateral earth pressure coefficient (K) is a variable which accounts for many factors, and depends on pipe diameter, diametric reduction due to Poisson's effect, soil dilation, etc. As noted by F i l z and Duncan (1996), K for loose sand may be close to at-rest conditions, and may be greater than 1 for compacted sand. Karimian (2006), based on the measured soil pressures on steel pipes, indicated that the initial value of K wi l l be close to at-rest conditions and would increase during axial pipe pullout (likely as a result of shear-induced dilation at the interface constrained by surrounding soil). With this background, the value of K for use in the proposed closed-form solution was back-calculated to match the peak measured frictional resistance. A value of 8 = 23.4° as 91 per Section 4.1.3 was used for this calculation. A s shown, in Figure 4.12, the back calculated K values for 60-mm and 114-mm pipes were 1.6 and 1.4. The following sections present a discussion on some of the difficulties in determining the proper value o f K . 1 I o (p = 39°, K = 1.4 (114-mm pipe) — 114D 60D — • 114mm-ASCE — -eOmm-ASCE (j) = 39°, K = 1.6 (60-mm pipe) 100 200 300 400 SP1 Displacement, x (mm) 500 600 Figure 4.12 Comparison of predicted axial pullout resistance from A S C E (1984) for 60-mm and 114-mm pipes. Actua l pressure dist r ibut ion around the pipe In real-life installations, the average soil pressure exerted on a buried pipe wi l l be different from that calculated considering an idealized average normal pressure (o n) given by Equation 4.12. (Note: This normal pressure is assumed in Equation 2.1 in deriving the axial frictional resistance). 92 2 Based on direct soil pressure measurements, Karimian (2006) pointed out that the actual pressure distribution around a buried pipe would be different from the above idealized pressure; e.g. in areas close to the pipe invert, smaller pressures were observed due to difficulties in soil compaction in that location. The effect of soil dilation As pointed out by Karimian (2006), the value of K wi l l vary significantly during axial pipe pullout mainly due to dilation of the sand, around the pipe. Although, the results from direct shear tests indicated no significant dilation at the MDPE-sand interface (O'Rourke, 1990), it is reasonable to assume that the soil immediately around the pipe would tend to dilate due to shearing action during an axial pipe pullout. The resulting volumetric expansion is contained by the surrounding soil mass. Accordingly, the dilation would increase the normal force on the pipe and, in turn increase the frictional resistance. If Equation 2.1 is to be used, this effect could only be accounted with an increased K value, since there is no other parameter to account for this effect. The numerical analysis conducted by Karimian (2006) further supported this observation. In some axial pullout tests by Karimian (2006), strips of colored sand were placed, adjacent to the pipe surface (see Figure 4.13) to observe the thickness of active shearing 93 zone around the pipe during axial pullout tests. When pipe was exhumed after testing, the thickness of the active shearing zone was found to be ~ 2.5 mm for roughened steel pipe-soil interface. However, when colored sand was placed around a concrete-filled P E pipe (i.e. " r ig id" P E pipe), the thickness of the affected zone was measured to be somewhat smaller, i.e. 1.6- 2.0 mm. It is possible that this may be due to smaller interface friction angle between P E and sand, resulting a smaller dilation around the pipe. Based on these limited observations, it may be possible to hypothesize that the increase in the value of K during axial pullout for buried P E pipe would be lower than those for roughened steel pipes. The effect of diametr ic changes in flexible pipes From numerical analysis, Karimian (2006) also showed that K depends on the pipe diameter (see Figure 4.14), resulting larger K values for smaller diameter pipes. This would, perhaps, explain the relatively larger K value obtained in 60-mm pipe compared to 114-mm pipe when the value of K was estimated using the same 8 value (see Figure 4.12). In flexible pipes, the diametric reduction due to increase in axial force (due to Poisson's effect) would allow volumetric expansion of soil in the shear zone during the pipe pullout process. Thus, this should reduce the normal pressure on the pipe. Therefore, the resulting overall K value could be lowered in the opposing manner to the likely increase in K due to dilation. These effects would be further complicated due to presence of 94 pressurized natural gas in the pipe. Clearly, there is a need to undertake further research to study the variation of K during relative axial soil movement and their effects on the pipe-soil interaction in P E pipes. In addition to the reduced dilation (thus smaller K value), surrounding soil could form a soil arch around the pipe with the diametric reduction of the pipe. This would prevent the entire weight of the soil column acting on the pipe (Spangler and Handy, 1982). This reduction in normal stress would result in smaller friction, in turn increasing the mobilized frictional length (progressive failure effect). Eventhough it is difficult to make a clear distinct between arching effect and reduction of soil dilation, for simplicity, the combine effect could be considered to be accounted through variation of K. 95 Figure 4.13 Affected shearing zone in P E pipes, detected from the disturbances in the .strips of colored sand (Karimian, 2006) 96 ~- 2.5 4 3.5 4 3 + ! 4 0.5 4 <p = 45° Ej « 94 (O j /P , ) " Mpa (i 0.5 I 1.5 Effect of dilation; Kxpansion at pipe surface (mm) 2 Figure 4.14 Variation of K with pipe diameter (Karimian, 2006) As the soil dilation and pipe diametric reduction depend on the shear strain of the soil, a single value for K would not represent the entire shear characteristics at the pipe-soil interface. In fact, K w i l l decrease with increasing shear resistance as a result of reduced dilation and diametric reduction of pipe. 4.1.4 Comparison of axial pullout resistance with other guidelines The pullout resistance can be calculated from several other methods proposed in literature. Such methods proposed by Danish Submarine Pipeline guidelines (1985) and McAl l ister (2001) can be presented using Equations 4.12 and 4-13, respectively. 97 4Wp (2 + K) -*—(2 + K) tan<p [4-13] 2 r v 2 3 n Tu= 2Dy(H--) + Wp tan<p [4-14] Note: W p is the weight of the pipe and considered to be negligible for these testing conditions. Using the same material properties (K = 1.6 and O =39°) as used in the A S C E (1984), the values of Tu can be predicted using Equations 4.13 and 4.14 as in Figure 4.15. As noted, these methods showed significant deviations from the measured pullout resistance of the P E pipes. 98 Figure 4.15 Prediction of the axial soil resistance using Danish Submarine Guidelines (1985) and McAl l is ter (2001). 4.1.5 Residual pul lout resistance As observed in steel pipes, at large displacements, a relatively loose soil structure is expected to be formed around the pipe (Karimian, 2006). Using a large strain friction angle of 33° (the constant volume friction angle) and K of 0.5, corresponding to loose soil conditions, the pullout resistance for pipe diameters 114-mm and 60-mm were calculated, and the results are presented in Figure 4.16. A s may be noted, these computed values are in good agreement with corresponding axial force obtained from Test No . 114Dand 60D. 99 8 1 -0 -|—• 1 1 1 1 1 1 L _ J 1 1 1 1 1 | _ J 1 1 1 1 1 , , , 1 , , 1 L _ _ 0 100 200 300 400 500 600 SP1 Displacement, x (mm) Figure 4.16 Predicted residual frictional resistance from loose soil properties ((p = 33° , K=0.5) 4.1.6 Response of straight pipe subject to reverse loading In Test No 114D, after completion of axial pullout loading in the initial direction, the pipe was pushed back in the opposite direction. For the considered pipe length, 114-mm diameter pipe is very unlikely to buckle because of the applied compressive load. The force-displacement characteristics are shown in Figure 4.17 for both the loading directions. The mobil ized frictional force in the pushing direction was larger than the first pulling force. As observed in Section 4.1.5 and as also mentioned by Stewart et al (1999) and Karimian (2006), after the first loading increment, a loose soil structure wi l l be 100 formed around the pipe. However, in this test, due to diametric expansion of the pipe in turn, w i l l increase the normal stress and the pullout resistance. Figure 4.17 Force-displacement relationship for the pull ing (positive) and pushing (negative) directions. 4.2 Lateral pipe pullout tests Two lateral pipe pullout tests (Test No . 114D-(2.5m) and 114D-(1.5m)) were performed in dense sands with 114-mm pipes. These tests were undertaken in view of developing an understanding of the lateral loading response and, in turn, to generate a data set for potential future numerical modeling. The pipes were also instrumented with strain gauges and pipe displacements using string potentiometers were measured at selected locations. 101 Two strain gauges were placed at a selected cross section, in diametrically opposite locations of the bending plane of the pipe (see Figure 4.18). Strain gauges «-->r-! Loading direction \ \ ! Bending plane Figure 4.18 Placement of strain gauges in opposite ends of the bending plane, for lateral pullout tests. 4.2.1 Observed load-displacement-strain response for lateral loading The pipe configuration of Test No 114D-(2.5m) test is shown in Figure 4.19 and the remaining test configurations are given in Appendix B. The pullout resistance measured (from one of the load cells) for a 2.5-m and 1.5-m long pipes are shown in Figure 4.20. A separate pullout test was performed to find the frictional resistance of the high strength steel cables connecting the pipe to the actuator head. The pullout resistances of the pipe were corrected accordingly. 102 SP1 Figure 4.19 Test configuration for Test No. 114D-(2.5m). The strain gauge readings allowed calculating the bending moments and tensile forces resulting from lateral movement, assuming that strains are small enough to use the simple bending formula. The strain gauge readings at the diametrically opposite location were almost identical in magnitude but opposite in direction (one in tension and other in compression). This confirmed that the pipe was mainly under bending for the selected length of the pipe. 103 14 0 - | L _ J — , — , — , — | — , — , — i — , — | — i — i — , — i — | — i — i — i — ( — | — i — i — . — i — f — 0 100 200 300 400 500 SP1 Displacement, x (mm) Figure 4.20 Lateral load vs displacement relationship of Test No . 114D-(2.5m) and 114D-(1.5m). The bending moment distributions computed for different end displacement levels during Test No. 114D-(2.5m) are shown in Figure 4.21. This also demonstrates the typically observed movement of the location of maximum bending moment along the pipe with increasing displacements (the location of maximum moment was at the pipe end at the start and moves towards the mid-length). 104 80 Length along the branch pipe (mm) Figure 4.21 Bending moment distribution of along the length of the pipe in 114D-(2.5m) test, for different displacements of the pipe ends. The typical active and passive soil wedges were observed in both lateral pullout tests. These three-dimensional surface deformations were not mapped because of their complex surface profiles. Note that in steel pipes, the active and passive soil wedges could be traced easily in two-dimensions, as the pipe behaves as a rigid object. A s a result of bending of the P E pipe in the horizontal plane, larger soil deformations (soil accumulations) were observed closer to the middle of the pipe length and smaller surface deformations at the ends. 105 4.3 Summary In analyzing the buried P E pipe systems, it is imperative to understand the basics of pipe-soil interaction occurring in P E pipes in order to understand the performance of more complicated pipe configurations. With this in mind, four full-scale axial pullout tests and two lateral pullout tests were performed to obtain the necessary modeling parameters. In analyzing the axial pullout tests, a closed-from solution was derived, taking into account the nonlinear stress-strain behavior of P E pipes. From this closed form-solution, the strain, force and mobil ized frictional resistance of the pipe could be obtained, for a known relative ground displacement. The ultimate mobil ized frictional length and the corresponding relative displacement at failure could also be predicted from the formulation. Further studies are needed to understand the interaction response, especially with regard to K and 5, to obtain more accurate modeling parameters for this method. In the reverse loading test, larger pullout resistances were observed during the second loading increment along the pushing direction, contrary to the typical observations made on steel pipes or in stiff P E pipes. The increase in pullout resistance was judged to be a result larger K value arising from diametric increase of the pipe. Two lateral pullout tests were performed to obtain the necessary data for potential future numerical modeling. The strain gauges attached to the pipe were used in deriving the bending moment and tensile force distribution along the length of the pipe. 106 CHAPTER FIVE FULL-SCALE PULLOUT TESTS ON B R A N C H PIPE SYSTEMS The distribution pipe networks essentially consist of complex configurations formed from various pipe components (e.g. trunk pipes (main), branch pipes (lateral), tees, bends, etc.). As the overall response of the pipe network is complex and difficult to analyze, it is important to understand the performance of such pipe component using full-scale model tests. Such model tests would not only enable identifying the key parameters (boundary conditions, soil conditions, pipe components, etc.) that would affect the overall response of the system, but they also would allow recognizing the l ikely failure locations and failure modes in these pipe configurations, under different testing conditions. The test results of eight branch pipe tests conducted to investigate the performance of the following three pipe configurations presented in this chapter. The test conducted were: 1. 60 mm trunk pipe and 16 mm branch pipes (Tests nos. 60D16-*) 2. 114 mm trunk pipe and 16 mm branch pipe (Test nos. 114D16-*) 3. 114 mm trunk pipe and 42 mm branch pipe (Test nos. 114*42-*) The test results include pullout resistance (load), displacement and strains at selected locations of these three branch pipe configurations. This chapter wi l l also discuss the 107 analysis of these pipe configurations with regard to selected slide geometry and boundary conditions replicated inside the soil box. 5.1 60-mm trunk pipes with 16-mm branch pipes (Test nos. 60D16-U and 60D16-R) The pipe configuration for the two Tests (Test Nos. 60D16-U and 60D16-R), along with the strain gauges and string potentiometers is shown in Figure 5.1. Pulling direction SP1 -4-a rO SP2 SP3 Load Cell : r=nftn SG1 SG2 1.0 0.75 >< > Branch Pipe'' S G 3 SG4 S G 5 0.75 0.75 < X > f9 fi S G 6 A fi Trunk Pipe SP4 Load Cells or string potentiometers Figure 5.1 Pipe configuration for Test No . 60D16-* and 114D16-* with strain gauge (SG) and string potentiometer (SP) locations in trunk pipe. 108 The butt outlet tapping tees were employed at the trunk-branch pipe connection. In all tests, at small displacements (<16 to 20 mm), the strain readings showed similar strain distribution to that of a pull ing of a straight pipe. However, as the displacement of the leading end exceeds 20 mm (approx), the strain in front of the tee increased relative to the leading end of the pipe (see Figure. 5.4). In every test, the strain gauges were attached to the crown of the pipe. In some tests, additional gauges were mounted to the pipe's invert to examine the uniformity in strain distribution at a cross section, especially in the pipe section leading up to the tapping tee. After a certain displacement, strain gauges attached to the crown showed larger strains than at the invert (see Figure 5.2). This localized strain in the trunk pipe was caused by the anchoring of the tapping tee. Immediately behind the tee, the measured strains were very small or even become negative (compressive strains) at large displacements. 109 10000 -2000 - 1 1 SP1 Displacement (mm) Figure 5.2 Strain gauge readings (at crown and invert of the pipe) at a location of 0.75 m from the leading end of the pipe in 114L42-R test. The response of the configuration observed using the string potentiometers attached to the trunk pipe is shown in Figure 5.3. As may be noted, the tees experienced very small displacement even after the leading end had experienced displacement in order of 580 mm. Clearly the presence of these tapping tees have caused significant amount of anchoring of the trunk pipe at the tee. Accordingly, the trunk pipe would experience very large strains leading up to the first tee. This was further confirmed by the large strains recorded in SGI and SG2 strain gauges (see Figure 5.4). As a consequence of this, the branch pipe would endure relatively small deformations. As a result, no detectable movement of the branch pipe was observed at a distance of 1.6-m (buried length of branch pipe, inside the soil box) away from the tee. In other words, in Test Nos. 60D16-U and 60D16-R, no displacement or 110 load was recorded in the string potentiometers or load cells placed at the end of the branch pipes, (i.e. points " X " and " Y " in Figure 5.1). It is expected that, as the boundary conditions of these branch pipes have no impact, the test results should be identical for these two tests. The load-displacement characteristics for two tests plotted in Figure 5.5 are essentially identical confirming this expectation. The strain distributions along the trunk pipe for the two tests were also found to be almost similar, again supporting the above. However, it is likely that at relatively large displacements of the trunk pipe (beyond 0.6 m of displacement), the end boundary condition of the branch pipe could contribute to the pullout resistance. A t this level of displacement, the performance of the pipe system wi l l depend on the amount of restraint provided at the branch pipe end. 600 0 100 200 300 400 500 600 SP1 Displacement (mm) Figure 5.3 String potentiometer readings at different locations of the pipe in 60D16-U 111 80000 60000 40000 w 2. ™20000 CO -20000 Edge of the box - * -2 mm mm -*-10 mm • 20 mm - * - 50 mm -•-100 mm 500 1000 Edge of the box-^-1500\. 2000 First tee 2500 3006^ 3500 Second end Length along the pipe (mm) Figure 5.4 Strain distribution along the trunk pipe for different leading end displacements in 60D16-U test. 14 6 0 D 1 6 - U • 6 0 D 1 6 - R 100 200 300 400 SP1 Displacement, x (mm) 500 600 Figure 5.5 Load -displacement characteristics for the 60D16-U and 60D16-R tests. 112 5.2 114-mm trunk pipes with 16-mm branch pipes (Tests Nos. 114D16-U and 114D16-R) The test configuration and the type of tapping tees used for these two tests were similar to that of 60D16-* tests. The pullout resistance characteristics (i.e. with respect to leading end displacements) for Test Nos. 114D16-U and 114D16-R are shown in Figure 5.6. The strain distribution along the length of the trunk pipe for Test No. 114D16-R is shown in Figure 5.7. The 114-mm diameter trunk pipe exhibited smaller strains in comparison to the tests with 60-mm trunk pipes discussed in Section 5.1. A s may be noted, the strain distribution is not significantly different from the unrestrained test (114D16-U). It appears that the smaller stiffness branch pipe (i.e. 16-mm pipe) is not capable of creating significant increase in resistance/strain in the trunk pipe, even i f restrained. Although, the strains in the trunk pipe were small (max ~ 2.0% measured), the strain gauge readings and string potentiometers attached to the branch pipes indicated considerable amount of deformation of the branch pipes, even after a slight displacement of the trunk pipe, (i.e. strains measured in branch pipes were greater than 3% at an displacement of about 150 mm of the leading end of the pipe). This signifies that 16-mm branch pipes are greatly susceptible to failure when connected to a trunk pipe with large stiffness. 113 Figure 5.6. Force-displacement characteristics of 114D16-R and 114D16-U tests. 20000 -10000 Length along the pipe (m) Figure 5.7. Strain distribution along the 114-mm pipe for different displacements of the leading end (Test No . 114D16-R) 114 Although the performance of the trunk pipe in Tests No . 114D16-U and No. 114D16-R were similar, the branch pipes in these two tests exhibited significantly different modes of failure. Visual observations made on specimens upon completion of testing (and removal of soil overburden) provided an opportunity to appreciate these differences (see Figure 5.8 and 5.9). The branch pipe deformations were found to be confined to a limited region (extending to a length of about -0.3 m within the trunk pipe axis, as per observations made on exhumed pipe after testing, see Figure 5.10). When restrained, the 16-mm branch pipes showed considerable amount of yielding before eventually separating at the tee (see Figure 5.8). In the unrestrained test, inspection o f pipe after the test showed buckling of the branch pipe immediately next to the tapping tee (see Figure 5.9), thereby possibly reducing the pipe cross section available for gas flow. Since there is no restraint at the other end, and once a plastic hinge has formed, the branch pipe seems to have dragged along the trunk pipe without any significant increase in strain in the branch pipe. Nevertheless, it should be noted that the frictional resistance to axial movement is very small (~ 0.24 kN/m) in 16 mm diameter pipes (Anderson, 2004). These observations are also supported by previous preliminary research observations reported by Anderson (2004). 115 Figure 5.8. Yielding induced failure of the branch pipe at the tee, in Test N o . l 14D16-R. Figure 5.9. Buckl ing induced failure of the branch pipe at the internal stiffener of the tee, in Test No . 114D16-U. 116 Figure 5.10. Deformation of the branch pipe and movement of the P V C protective sleeve along the pipe in Test N o . 114D16-U (Anderson, 2004). 5.3 114-mm t runk pipes wi th 42-mm branch pipes (Tests No. 114*42-*) Tee-connections joining 114-mm trunk pipe and 42-mm branch pipe were also tested in this study. A typical pipe configuration is shown in Figure 5.11. In contrast to the two branch pipe test configurations discussed above, in this pipe configuration, socket outlet tapping tees were used in place of butt outlet tapping tees (shown in Figure A . 1 (b) of Appendix A) . The tests were performed both in loose and dense sands. The load-displacement characteristics for loose and dense sand are shown in Figure 5.12 and 5.13, respectively. The corresponding measured load cell reading at the end of the branch pipe for the restrained configuration are also given in the same figures. In each figure, the load difference between "restrained" and "unrestrained" curves directly reflects the effect of 117 branch pipe constraint on the trunk pipe. As may be noted from the figures, the load difference observed for the configurations in dense sand were similar to that observed for the configuration in loose sand. It is also of importance to note that, this difference is also similar to the increase in resistance at the branch pipe end. These observations suggests that, irrespective of the soil density, after mobil izing the restraint condition at the pipe end, much of the increase in the pullout resistance is contributed from the fixity of the branch pipe, when branch pipes are restrained from movement. SP1 7 HI Load cell SP2 ^ < 1 ^ 1.0 SG1 SG2 SG3 LT SG4 n n 1.0 1.0 X — SG5 & 6 0.5 cbji «Y SG6 & 7 42 mm Branch Pipe String potentiometer or ». Load cell • SP3 114 mm pipe Figure 5.11. Test configurations for 114*42-* tests showing strain gauge (SG) locations along the trunk pipe. 118 14 SP1 Displacement, x (mm) Figure 5.12. Load-displacement characteristics of 114L42-U and 114L42-R tests. Figure 5.13. Load-displacement characteristics of 114D42-U and 114D42-R tests. 119 In all four tests conducted with this configuration, the failure occurred at the tee, where the cross section was smaller than the diameter of the adjoining branch pipe (see Figure 5.14 and A . l in Appendix A ) . In 114D42-R test, the branch pipe completely separated from the tee, causing a drop in pullout resistance (see Figure 5.13). A lso in other tests, a similar fracture was observed but with varying degrees of damage. A s expected, in all restrained tests, the observed fracture was wider than in unrestrained test. Considering the location of failure, it can be deduced that, the damage was greater in dense tests than in loose tests. For this pipe configuration, the maximum pullout resistance recorded at the leading end (or the capacity of the pipe system) would depend on the capacity of the socket outlet tapping tee. Unfortunately, since the pipes were not pressurized due to safety reasons, the actual failure point was impossible to identify. However, by examining the degree of damage in each test, an approximate maximum allowable axial displacement of the trunk pipe could be established, i f needed. For example, based on the test conducted, the allowable displacement is judged to be around 550-mm when the branch pipe is unrestrained, and placed in loose sand. Considering the variability in field installation conditions, such qualitative values would be sufficient to make a rational judgment about the pipe's performance. 120 Figure 5.14. Formation of plastic crack at the tee in 114D42-U test As also observed in other two pipe configurations (i.e. Test series 114D16-* and 60D16-*), surface deformations were not observed even with this pipe configuration with large diameter branch pipe. Nevertheless, surface deformations were observed in the test with a butt-tee connection (i.e Test No . 114D42 (butt tee)). The test results o f this butt-tee connection test are not discussed in this section but the results are presented in the Appendix C . In the current series o f testing, all tests were performed at a constant burial depth of 0.6 m. A t this burial depth, the lateral movement of the branch pipe would result in deep bearing failure as shown in Figure 5.15(a) and the resulting deformations are likely to confine to an area surrounding the branch pipe. These deformations are not 121 l ikely to bring about large slip surfaces, which are visible at the surface. However, Anderson (2004) observed surface deformations in similar branch pipe tests. This perhaps attributable to the smaller burial depth in some of the pullout tests which in turn could result in these typical passive and active soil wedges which can be observed at the ground level (see Figure 5.15 (b)). Figure 5.15 (a) Punching failure at deep burial depths (b) Typical formation of passive and active soil wedges at shallow burial depths, when pipe moves laterally. 122 5.3.1 Induced rotation in the trunk pipe due to the tapping tee The rotation of the trunk pipe was also monitored in these four tests. This was accompanied by tracking of the crown of the trunk pipe (at the leading end). The observations suggested that, the trunk pipe would rotate when the branch pipe was restrained against movement (see Figure 5.16). The rotation was observed after a displacement of -175 mm of the leading end and then continued at a rate of -0.39 rad/m. In tapping tee connections, the branch pipe is connected with an eccentricity (e) to the axis of the trunk line (see Figure 5.17), compared to a typical butt tee connection, where the pipe axes wi l l be in the same horizontal plane. As a result of this eccentricity, the joint can induce torque/rotation in the trunk pipe. The pipe coupling used at the leading end of the pipe was capable of accommodating the rotation. The excavated pipe did not show any form of deformation to suggest twisting of the trunk pipe. It was likely that the resulting deformation was borne by the smaller cross section of the tee. This rotation would also explain why the puncture occurred at the underside of this cross section. The amount of observed rotation was approximately similar in loose and dense tests. It appears that, this kind of rotation is likely to occur only in restrained large diameter branch pipes, in which the transmitted eccentric load is considerable. It is important to note that, no rotation of the trunk pipe was observed in for the 42-mm branch pipe with no restraint (e.g. Test Nos. 114*42-U), and also in branch pipes with smaller diameter (e.g. Test Nos. 114D16-* ) regardless of the restraint condition or in 114*42-U tests. 123 Figure 5.16. Rotation of the trunk pipe in 114D42-R test. Figure 5.17. Torque induced in the trunk pipe due to the eccentricity (e) of the loading axes. 124 The strain distribution for the trunk pipe in 114D42-R test is shown in Figure 5.18. When the trunk pipe is considered, despite the large diameter and stiffness of the branch pipe, it appears that the flexibil ity at the small cross section at the tee (see point " X " in Figure 5.17) would prevent transfer of any significant stresses from the branch pipe to the trunk pipe. For instance, in 114D42-R test, the maximum strain recorded closer to the tee was only 1% even at a displacement of 100 mm at the front end (see Figure 5.18). This measured strain was not significantly different to that of Test No . 114D42(tee only), (i.e. the test in which the branch pipe was removed and pullout test was performed only with a tapping tee attached to the trunk pipe, see Appendix C). It was also observed that even after the failure of the tee, the branch pipe was in good condition. 10000 -2ooo 1 ! •—'- ; 1 1 — i Length along the pipe (m) Figure 5.18. Strain distribution along the trunk pipe for different displacements of the leading end, in 114D42-R test. 125 To summarize, Table 5-1 shows the likely or observed failure modes in eight branch pipes tests performed. A s may be noted and previously described, these modes and location of the failure are influenced by relative pipe sizes, type of tapping tees, slide geometry, amount of ground displacement, mobil ized frictional length along the trunk pipe, soil density and end boundary condition o f the branch pipe. These results for full-scale tests provide insight into the potential failure mechanisms under real-life conditions. Table 5-1 L ike ly or observed failure modes and locations for the considered pipe configurations. Test no. Observed failure modes 114D42-U Rupture/ cracking observed at the tee to varying degree. No damage was observed in branch and trunk pipe. 114D42-R 114L42-U 114L42-R 114D16-U Branch pipe failed immediately after tee from buckling 114D16-R Branch pipe failed due to Yield ing and necking 60D16-U Very large strains observed in trunk line 60D16-R 126 5.4 Analysis of branch pipe configurations A n analysis procedure for branch pipes subjected to relative ground movements in the direction of trunk pipe alignment is discussed in this section. The pipe-soil interaction occurring at the joint is very complex and the selected boundary conditions, pipe sizes, soil and material properties, type of joint and slide geometry can affect the overall capacity of the pipe system to undergo permanent ground displacements. One o f the approaches that could be used to analyze such complicated systems would be to consider the system as an assembly of discrete components. In this study, attempt was made to analytically capture the response of these components on an individual basis. The response of the full system would then be obtained by combining the individual components. Such attempt was made in the following analysis, for the selected pipe configurations. To analyze the pipe system, at first, it is essential to identify the vulnerable sections of the pipe system and the associated failure mode. 5.4.1 Identification of components of branch pipe system for analysis. Using observations from the pullout tests, typical axial strain distribution along the trunk pipe can be developed for a system with branch pipes, when the trunk pipe is subjected to axial movement. A s noted in this type of pipe configuration, the critical strains are expected at two cross-sectional locations of the trunk pipe depending on the mobilized frictional length (xf) along the trunk pipe. 127 If the mobilized frictional length is small, as in case (i) of Figure 5.19, it is possible to experience larger localized strains on the trunk pipe closer to the tee, as opposed to the leading end of the pipe (see Section 5.1 to 5.3 for results from pullout tests). A s shown later, once full resistance of the tee is mobilized, the axial strains in the trunk pipe in front of the tee would not l ikely to increase appreciably any further with pipe displacement. Any increase in resistance/strain thereafter is associated with the increase in mobilized friction length beyond the tee. Furthermore, the localized strain component due to anchoring resistance of the tee would also likely to increase with increasing mobilized frictional resistance due to non-linear material behavior of the P E pipe, even after full resistance of the tee is mobil ized. In case (ii), i f the total mobil ized frictional length (xf) is large, the largest strains could be observed at the leading end of the pipe due to mobilized frictional length along the trunk pipe and anchoring forces of the tapping tee. Based on the above, the cross-sections B and A ' could be considered as the critical pipe sections in case (i) and (ii), respectively. However, the l ikely performance limits could be different at these pipe cross sections, depending on the anticipated failure mode, e.g. pipe could fail in yielding at A ' , whilst at B, the brittle-like failures or even yielding could occur. 128 Pipe under axial movement B Strain distribution A ' Pipe under axial movement B' Strain distribution (ii) Figure 5.19. Crit ical sections of the pipe for different total mobil ized friction lengths (x f) along the trunk pipe 129 When the leading end of the pipe (A) is considered, the total axial resistance (F T) on the trunk pipe is contributed from three distinct components as expressed using Equation 5.1. Fr=Ft+Fl+Fb [5-1] where, F, = Anchoring resistance of the tapping tee. F, = Frictional resistance caused by the axial movement of the trunk pipe. Fh = Resistance due to movement/anchoring of the branch pipe. These components were evaluated from the standpoint of contributing to the assessment of pipe performance in branch pipe configurations subjected to ground movement. 5.4.2 Anchoring resistance of the tapping tee (Ft) Considering the complexities of the soil interaction at the joint, it was decided to represent the response of the tee using a force-displacement (soil-spring) behavior. While the results from axial tests conducted on straight trunk pipes and branch pipes were available, it was noted that the response from an axial load test on a trunk pipe that has only a tee (but no branch pipes) would allow a more reliable estimation of the effect of the tee. With this in mind, a Test No . 114D42 was performed on a trunk pipe with a tapping tee (by removing the branch pipe from the tee). As observed in Figure 5.20, a localized increase in resistance between 120 mm to 220 mm of displacement was noted 130 during pullout test, in Test No . 114D-1. It is believed that this is due to "dragging" of the strain gauge wires. The anchoring resistance of the tee (F t) could then be obtained from the difference between the pullout resistance of Test No . 114D-1 on straight pipe and that from 114D42 (tee only), as in Figure 5-20. This approximation was considered to be valid as the strains in the 114-mm pipes were not significantly different between Tests No . 114D-1 and 114D42(tee only), thus the "shaft" frictional resistance would be similar. The resistance of the tee estimated as per above was found to be around 1.9 k N (see Figure 5.20). 12 o 3 4 0_ 24 •114D •114D42(tee only) 100 200 300 400 SP Displacement, x (mm) 500 Figure 5.20. Load-displacement characteristics of 114D and 114D42 (tee only) test. 131 Since the resistance is arising due to relative lateral movement against the soil in front of the tee, it was considered of interest to investigate the possibility o f using lateral bearing capacity factors available for pipes and vertical anchor plates for computing soil resistance of the tee. In this regards, the lateral resistance against the movement of the tee was computed using the Equation 5.2 below as per A S C E (1984). FT={NqyH)A [5-2] In this equation, A t is the projected area of the tee in a plane perpendicular to the direction of the movement. Using this approach, the computed value of F t is 2.1 k N , and it is in good agreement with the maximum pullout resistance o f the tee obtained from experimental work. Based on this outcome, it is judged reasonable to estimate the value of F t using Equation 5.2. It is noted that the value of N q for use in Equation 5.2 can be found from the respective H/D ratio and peak friction angle (see Figure 5.21), according to guidelines stipulated in A S C E (1984). In developing soil-spring models, addition to the value of maximum lateral resistance (F t), it is also required to determine the displacement (yh) required to mobil ize the value of F t . Since the test data from pipe testing did not provide reliable means of obtaining the value for yh, it was decided to estimate yh assuming yh/H = 0.02 (H = 0.6 m) in accord with the experimental observations reported for pipes and vertical anchor plates, in dense sand ( A S C E , 1984). Accordingly, for the tested pipes, it is estimated that the full 132 resistance of the tapping tee would be mobilized at a displacement of - 1 2 mm. Considering the above, it was determined that the interaction of the tee with the soil could be characterized with a bilinear representation given by a maximum anchoring resistance of 2.1 k N and yh of 12 mm (see Figure 5.22) for the considered burial depth and friction angle. Figure 5.21. Normalized lateral bearing capacity factors for different H/D ratios and friction angles ( A S C E , 1984). 133 2.5 V 2 o c JS - 1 5 CD or +-> I 1 D_ 0.5 4 0 ' F t = N q Y H A t «< • l • « 1 • l • . • 1 - • - i ~ : ' a 1 - : 1 • i • t J • i • i . : • * i I i • i - • : • ••> a 1 • | : 1 y h = 0.0'2H ? 1 i • - 1 ' ' 1 — 1 i 1 1 — 1 I 1 ' 1 1 i 1 1 1 1 i 1 1 1 1 100 200 300 Tee Dsplacement (mm) 400 500 Figure 5.22. Soi l spring characteristics for the tee at a burial depth of 0.6 m and in dense sand. To assess the validity of this approach, the above soil-spring model of the tee was then used in estimating the pullout resistance corresponding to another test conducted with only the tee connected to the trunk pipe, (.i.e. Test no. 60D16). The pipe configuration of this test was similar to the test series 60D16-*, except for the absence of branch pipes off the tee. Knowing the displacement of the tees (from string potentiometers attached to the tees), the anchoring resistance of these tees could be calculated from the soil-spring characteristics given in Figure 5.22. A lso the shaft friction of the trunk pipe was assumed to be not significantly different from that obtained from Test No . 60D-1 which is the test performed on 60-mm straight pipe. The total resistance at the leading end could then be computed by 134 the summation of frictional resistance along the trunk pipe and anchoring resistance of the tee. The estimated and measured force-displacement characteristics are shown in Figure 5.23, and the reasonably good agreement between these two curves confirms the suitability of using this force-displacement characteristic for representing the effect of the tee. 600 SP1 Displacement, x (mm) Figure 5.23. Measured and calculated (from soil spring behavior) pullout resistances of the 60D16 (tee only) test. Clearly, the soil-spring characteristic of the tee would depend on the burial depth, soil density and even on the type and size of tapping tee. It is proposed that, in the absence of direct experimental data, approximate soil-spring characteristics for the tapping tee in different soil and burial conditions are to be derived using N q values, stipulated for pipes or vertical anchor plates (e.g. A S C E , 1984). The corresponding y p value could also be 135 obtained from guidelines given for anchor plate and lateral pipe pullout characteristics (e.g. Audibert and Nyman, 1977 and A S C E , 1984). 5.4.3 Resistance due to movement/anchoring of the branch pipe ( F b ) The contribution of the branch pipe resistance (Fb) to the axial movement of the trunk pipe is a difficult parameter to estimate, due to several reasons. Firstly, typical branch pipes are small in diameter (e.g. 16 mm diameter); as such, any meaningful structural analysis is impossible because of low stiffness and strength of the pipes (Chapman, 1990). Furthermore, as mainly done in field situations, the 16-mm tubes are installed in a "snaked" manner (i.e. snaking), thus allowing greater flexibil ity when subjected to ground displacements. Although, due to these constraints, the structural analysis of the small diameter branch pipes in an exclusive manner was considered not practical, the impact of branch pipe (Fb) is considered in analyzing the trunk pipes. With this background, the fol lowing describes an attempt to represent the anchoring resistance of the branch pipe in form of soil-springs in analyzing the trunk pipe. 5.4.3.1 Resistance F b in 16-mm branch pipes Using Test No . 114D42 (tee only) and the derived soil-spring characteristics of the tee as per Section 5.4.2, the resistance offered by the tee and the frictional resistance of the trunk pipe (i.e. F t + F|) could be obtained. It can be argued then that the difference between above (F t + Fi) value and the results for the Test No . 114D16-U would yield the 136 resistance offered by the two branch pipes corresponding to this configuration. It is noted that, since the differential displacement between first and second tee in 114-mm pipe is minimal, it is reasonable to compute the respective force-displacement characteristic for one branch pipe by taking one-half of the total resistance computed for the two pipes as per above. It is also of significance to indicate that this derived force-displacement characteristic for the branch pipe would only be valid until the unrestrained end of the branch pipe begins to move (i.e. beyond this point, end restraint condition of the branch pipe would begin to affect the results). The force-displacement relationship computed based on experimental results as per above is shown in Figure 5.24. It is considered reasonable to assume that the displacement of the tee, and the movement of the branch pipe would be similar regardless of the diameter of the trunk pipe (for the diameters of trunk pipes typically used). If this assumption is valid, then the relationship in Figure 5.24 derived from 114-mm pipes should provide an opportunity to predict response of the Test series 60D16-*, and this was attempted as described below. Fb =0 .048x-0 .00025x 2 [5-3] Combining the results from Test No . 60D16 (tee only) and Figure 5.24 (Fb value), the total resistance (Fi + F t + Fb) at the leading end of the 60-mm pipe could be estimated. For computational convenience, the branch pipe contribution, Fb from Figure 5.24 was estimated using a fitted curve as given in Equation 5.3 which is also shown in Figure 137 5.24. The results show a good match between the predicted and the measured force-displacement characteristics of the 60D16-* tests. It is however noted that, in Equation 5.3 (where x in mm and Fb in kN) , the resistance offered by the branch pipe (Fb) should be limited to a value of -2 .5 k N , since beyond this point, 16-mm branch pipe would likely to fail from yielding. If this is not considered, the analysis o f the trunk pipe would become overly conservative with excessive branch pipe contribution. Computed from experiments Approximation 1.0 - - — -Displacement of the tee /(mm) Figure 5.24 Soil-spring characteristics for the unrestrained 16-mm pipe in dense sand 138 14 12 + — Calculated Measured 0 0 100 200 300 400 500 600 -2 SP1 Displacement/(mm) Figure 5.25. Measured force-displacement characteristics (at the leading end) from Test No. 60D16-U compared with those derived from soil-spring characteristics for tee and branch pipe. 5.4.3.2 Resistance Fb in 42-mm branch pipes The approach described in the previous section for 16-mm branch pipes were extended to the cases of 42-mm branch pipes as per below. In analyzing the 42-mm pipe, when restraint condition is mobil ized at the end of the branch pipe, the additional resistance measured in the trunk pipe (Fb) is mainly attributed to the resistance offered due to restraining effect of the branch pipe. This additional 139 resistance due to end restraint was observed to be similar in loose and dense sand (see Figure 5.12 and 5.13). Considering this aspect, branch pipe resistance for 42-mm pipe could be approximated by the linear relationship (see Equation 5.4) as shown in Figure 5.26. In Equation 5.4, units of Fb and x are k N and mm, respectively. Nevertheless the maximum resistance offered by the branch pipe wi l l be limited by the capacity of the socket outlet tapping tee. A s a result, a maximum resistance of 4.5 k N was imposed to avoid any overpredictions of the branch pipe contributions. Fb = 0.00114x [5-4] CD CL — 114D42-R —114L42-R F b = 4.5 kN y = 0.011 x 100 200 300 400 500 Displacement of tee after mobilizing end restrain, x (mm) Figure 5.26 Branch pipe resistance of 42-mm pipe when restraint condition at the end restraint condition is mobil ized. 140 The above soil-springs developed for branch pipe systems (Equation 5.3 and 5.4 could be used in determining the branch pipe contribution (Fb) of a trunk pipe under axial movement. However, it is evident that further experimental data is required to make a more confident analysis of the anchoring resistance offered by a branch pipe (Fb) when (i) 42-mm pipe is unrestrained and, (ii) 16-mm pipe is restrained at a smaller branch pipe length. Moreover, it is noted that this suggested soil-spring characteristics are only be valid for the stipulated soil and burial condition and it should be modified for other conditions. 5.5 Stress concentration due to tee and localized strains It is generally believed that the incidence of localized stress concentrations in plastic pipes would be less in comparison to steel pipes due to the higher ductility of the former. With respect to long-term performance, despite the better ductility, f ield inspections of pipe failures shows that most of P E pipe failures are attributable to brittle-type failures arising due to stress concentrations (NTSB, 1998) under general operating conditions (i.e. not necessarily from ground movement). In this context, it can be argued that localized stress concentrations that would occur due to ground movements at the tee-joints and their presence over long time duration may have significant effect on the integrity of the pipe system. When considering a location similar to " B " in Figure 5.20 (a location closer to the tapping tee), the analysis would become rather complicated due to these localized strains induced from anchoring of the tee and the branch pipes. 141 Figure 5.27 shows localized strain in the trunk pipe due to presence of: (a) tee and (b) tee and branch pipe together. As may be noted, the presence of the tee alone can cause a significant increase in strain in the trunk pipe. As observed in the 60D16-* test series, for a displacement of 100 mm at the leading end of the pipe the displacement of the tee was 25 mm; this would result in almost 8% localized strain in the trunk pipe (in front of the tee), whilst tee alone (i.e. Test No. 60D16(tee only)) can create a localized strain of 4.5% at the same displacement. Although it is not possible to capture the maximum localized strain from strain gauge readings within its measurable range, these strain gauge readings would still provide a good lower bound approximation of the local strain magnitudes. 80000 60000 c 40000 CO 20000 .•1B • i I : •° la f u A (b) With Branch pipe 1 , - 1 (a) Tee only (no branch pipes) 10 20 30 40 Displacement of the tee /(mm) 50 60 Figure 5.27. Localized strain (in front of the tee) in the 60-mm trunk pipe due to presence of (a) tee, and (b) tee and 16 mm branch pipe (i.e. Test series 60D16-* with branch pipes). 142 For the test conducted, these localized strains were much more important for the 60 mm pipe since the overall strain in the trunk pipe was considerable. Conversely, the localized strains measured in the 114 mm pipe were unlikely to risk the integrity of the trunk pipe. In other words, prior to the failure of the 114 mm trunk pipe, failure would be seen at the branch pipes (i.e. Test series 114D16-*) or at the tee (i.e. Test series 114*42-*). In these situations, there wi l l be warning signs (loss of gas supply to entities) prior to build up of large strains in the trunk pipe. It is also significant to note that the displacement of a trunk pipe at two locations along cannot be used to assess strains on a pipe. This is illustrated in Figure 5.28 using data for Test No. 60D16-U. A s may be noted, i f the axial strain in the pipe was estimated only using the measured displacement at the leading end and the tee (without considering the localized strains), those estimates would be much different from the strains observed from direct measurements. Such computed strains using differential displacements alone are not only unsatisfactory but also non-conservative. Therefore, unless strain gauges are employed as in these tests, these localized strains wi l l become difficult to capture. The conventional soil-spring analysis is also not capable of deriving these localized strains from numerical analysis. Instead, sophisticated continuous finite element methods could be employed to capture these strain concentrations. However, due to convergence problems arising from strain localizations and lack of calibration data, these analyses would also become tedious. As a result, more experimental research is needed to account these localized strain in a meaningful manner. 143 80000 60000 40000 w CO =20000 -20000 Leading end Trailing end-Strain gauge readings Strain calculated from differential displacements 500 1000 1500\ 2000 2500 3000^ .3500 First tee Second end Length along the pipe (mm) Figure 5.28 Measured and calculated strain distributions along the trunk pipe from strain gauges and string potentiometer measurements (at a leading end displacement of 100 mm) in Test No . 60D16-U. 5.6 Effect of rate of displacement on pullout resistance As noted in Section 2.1, very little is known about the rate dependency in axial pullout tests. In all tests conducted with dense soil conditions some limited data was acquired by changing the rate of pipe pullout displacement. For example, in Test No . 60D16-R, the rate of displacement was increased to 180 mm/hr (five times the initial speed) from the initial rate of 36 mm/hr after 460 mm of displacement, which resulted in an increase of pullout resistance of 8.6% (see Figure 5.5). In Test No . 114D42-U, a similar change of 144 the rate of pullout displacement caused an increase of pullout load by 5% (see Figure 5.13). In general, it was observed that the pullout resistance would increase with the increasing rate of displacement. In loose sand, no visible increase in resistance was observed, even at a displacement rate of 180 mm/hr. However, it is not known whether the increment in pullout resistance was merely due to increased stiffness in P E pipe or whether the soil properties also contributed to the additional resistance, even at dry state. In a viscoelastic material such as P E , the stiffness of the material would increase with increasing strain rate (higher modulus). At slower rates, the polymers wi l l have more time to orient their molecules to the loading direction, thus the resulting stress-strain characteristics wi l l be softer (see Section 3.1.2.2). The effect of rate dependency, however, was not investigated herein since this was beyond the scope of this current study. 5.7 Effect of soil density A s indicated in the test program (Table 3.4), a limited number o f tests were performed on pipes embedded in loose sand to assess the effect of backfi l l soil density on the performance of P E pipe configurations. In general, the observed soil resistance in dense sand was appreciably higher that that for tests conducted with loose sand. A s noted in Figure 5.29, in branch pipe configurations, the soil density appears to have a more significant effect when the branch pipe is 145 unrestrained than when it is restrained. Since the additional strain/resistance due to end restraint is considerable when pipe is restrained, the effect of soil density is not considerable. Hence, large increase in pullout resistance is not observed when pipe is branch pipe is restrained unlike in unrestrained case (see Figure 5.30). 16000 0 100 200 300 400 500 Displacement of the Tee/(mm) Figure 5.29 Strains (at a distance of 0.75 m from leading end) in loose and dense sand when branch pipes are unrestrained. 146 30000 -6-114L42-R 500 Displacement of the Tee/(mm) Figure 5.30 Strains (at a distance of 0.75 m from leading end) in loose and dense sand when branch pipes are restrained. 5.8 Impact of the P V C protective sleeve It should be noted that, thus far, there are no standard specifications for deciding the diameter, length and stiffness of the P V C protective sleeve used at typical tee connections as shown in Figure 5.10 (NTSB , 1998). The sleeve was intended to provide additional bending stiffness at the tee and protect branch pipe from shear forces N T S B (1998). As in field installations, all tests in this study were conducted with tee configuration including P V C protective sleeves. 147 It was observed that, especially in 16-mm diameter branch pipe connections, the inside diameter of the sleeve was slightly larger than the diameter of internal stiffener of the tee, thus the sleeve was not properly locked in place. Consequently, observations made on exhumed pipe after pullout testing indicated that the sleeve had moved along the branch pipe with increasing deformation of the branch pipe. This suggested that, a sleeve that is not properly attached to the tee (which is the current practice) would not significantly contribute to increasing the bending stiffness of the branch pipe as intended. It appears that a sleeve that is properly attached to the tee's internal stiffener would contribute to the bending stiffness of the branch pipe. This would become more important i f the anticipated displacement of the tee and the branch pipe is significant (when branch pipe is connected to a large stiffness trunk pipe). A lso in 114-42mm socket outlet tee connections, the sleeve would not affect to the overall performance, since the weaker section of the tee was not reinforced by the sleeve. 5.9 Summary Altogether eight full-scale pullout tests were conducted on three pipe configurations. For the considered pipe configurations and accompanied boundary conditions, the pipe failure modes and the critical locations were explored. 148 It was observed from these pullout test results that: • In smaller diameter trunk pipes (i.e. 60-mm), the trunk pipe would experience large strains mainly due to anchoring of the tapping tee and branch pipe. • In large diameter trunk pipes (i.e. 114-mm) with smaller diameter branch pipes (i.e. 16-mm), the branch pipes would become vulnerable even under a small ground displacement. In these situations, P V C protective sleeve that is properly locked in place could increase bending capacity of the branch pipe. • In large diameter trunk pipes (i.e. 114-mm) attached to a relatively large diameter branch pipe (42-mm), the capacity of the pipe system to undergo ground displacement was limited by the capacity of the socket outlet tapping tee. In other words, the failure was observed at the socket outlet tapping tee. After considering these failure modes and the critical locations of these three pipe configurations, it was decided to analyze only the trunk pipe, considering the difficulties in analyzing the branch pipe. Due to complexity even in this analysis, the impact of various components on trunk pipe is represented by experimentally derived soil-springs. The following gives a summary of such representations formulated to account the anchoring resistance of the tapping tee and branch pipe (both restrained and unrestrained). 149 The respective soil-spring characteristics of a tapping tee could be approximated by a bilinear relationship. The maximum resistance of the tee is derived using Equation 5.2, in which the lateral bearing capacity factor (N q ) is obtained from results presented for lateral movement of vertical anchor plates and pipes (e.g. A S C E , 1984). The corresponding displacement (yh) is obtained from the same guidelines. For numerical modeling, the contribution of the branch pipe (Fb) was derived for 16-mm unrestrained pipe and 42-mm restrained pipes as in Equation 5.3 and 5.4, respectively. 150 CHAPTER SIX S U M M A R Y AND CONCLUSIONS The performance of buried polyethylene (PE) natural gas pipeline systems in areas subjected to permanent ground displacements is a key engineering concern since geotechnical hazards can be a major cause of damage to these utilities. Although polyethylene has become the industry standard in the gas distribution industry, there is only limited experience with pipeline materials other than steel. A s a result lack of alternative approaches, pipe-soil interaction models used in steel pipes are often considered for analyzing P E pipe configurations as well. However, due to combined effects of relatively high pipe material flexibility and interaction aspects with soil (soil dilation, arching etc), these simplified models would not sufficiently represent the basic pipe-soil interaction occurring in P E pipes. Furthermore, the commonly occurring pipe connections in the distribution pipe networks render significant impact on the integrity of the pipe system. With this background, a detailed research program involving full-scale physical modeling of buried pipeline systems has been undertaken at the University of Brit ish Columbia (UBC) . As a part of this work, several full-scale pullout tests were performed on straight and branch pipe configurations, mainly with the objective of understanding the basics of pipe-soil interaction and the response of the branch pipe systems undergoing complex 151 interactions at the pipe connection. For the purpose of clarity, the summary of the test results and the key findings are presented under two separate topics of (i) axial and lateral pipe pullout tests, and (ii) pullout tests on branch pipe configurations. 6.1 Axial and lateral pipe pullout tests With the objective of understanding the fundaments of P E pipe-soil interaction, four straight pipe pullout test and two lateral pipe pullout tests were performed on P E pipes of diameter 60 mm and 114 mm. In axial loading situations, unlike in steel pipes, the pullout response of P E pipes as noted to be significantly affected by the material behavior of P E . Thus, in P E pipes, in addition to the parameters that are commonly considered in the assessment of steel pipe response (e.g. burial depth, soil density, etc), the flexibility of the pipe (diametric reduction due to Poisson's effect and elongation) and nonlinear stress-strain behavior (due to viscoelasticity) would also influence the interaction response of buried P E pipes. To account the nonlinearity in stress-strain behavior of P E , a closed-form solution was developed. This would provide the framework for the analysis of straight pipe sections undergoing permanent ground displacements. The solution was capable of capturing the response observed in the full-scale tests. This closed-form solution can be applied to field situations to estimate the force, strain and the mobilized frictional length along the pipe for a known relative axial ground displacement. The solution provides the opportunity to derive the maximum possible mobilized frictional length (x p) along the pipe and the 152 corresponding displacement at failure (due to yielding) of the pipe. This type of an approach is not only useful in evaluating the operational safety of a pipeline but also useful as a screening tool to identify the critical locations of a pipe system located in areas of relative ground movement. For example, i f a pipe connection (e.g. saddle fusion tapping tee) is located in the pipe at a distance greater than the maximum mobilized frictional length (x p) from a point where relative ground movement has been observed (i.e. ground crack), then the failure is likely to occur at the ground separation point. In other words, in this situation, no localized strains wi l l be generated closer to the tee, because, prior to mobil izing the anchoring resistance of the connection, the pipe would yield at the ground separation point) It is also noted that, in this closed-form solution providing the analytical framework, the accuracy of the predictions can be further improved with better understanding of the response of K (lateral earth pressure coefficient) and 8 (interface friction angle) for different levels of shear strain in the soil, i.e. the value of 8 depends on the level of shear strain in soil, whilst the value of K would also depend on the level of dilation/shear strain and the diametric reduction of the pipe. Based on initial evaluations, the value of K in P E pipes was deduced to be smaller than that of steel pipes, suggesting that this would possibly be due to diametric reduction in P E pipe. This hypothesis was further supported by data from a reverse loading test, in which a larger resistance was observed than the first loading increment when pipe is under compression, suggesting that this would have been a result of the increased normal pressure on the pipe from diametric increase of the P E pipe. 153 Two lateral tests were performed to obtain necessary modeling parameters for potential future calibration/validation of a numerical model. It was observed that reasonably accurate bending moment and the tensile force distribution in the pipe could be obtained from the strain gauges attached to the pipe. In the preliminary analysis of these test results showed that the point of maximum bending moment moves along the pipeline with increasing pipe displacement, confirming what has been reported in the literature. 6.2 Pullout tests on branch pipe configurations Since the gas distribution systems consist of complex pipe networks, it was considered important to test typical pipe configurations to recognize the impact of various pipe components. With this intent, eight branch pipe pullout tests were conducted on three pipe configurations with different pipe sizes, soil and boundary conditions. Commonly used saddle-fusioned tapping tees were used at the branch-trunk pipe connection. Based on the test observations, it was recognized that the most l ikely failure locations in a branch pipe system could be at the trunk pipe, branch pipe or even at the tapping tee, depending on test conditions (relative pipe sizes, type of tapping tees, amount of ground displacement, mobil ized frictional length along the trunk pipe, soil density and end boundary condition of the branch pipe). For example, in low stiffness trunk pipes, (e.g. 60-mm pipe), the anchoring resistance of the tee and the branch pipe is significant such that it would create large localized strains at the vicinity of the tee-connection. In trunk 154 pipes with relatively large stiffness (e.g. 114-mm pipe), generally the branch pipes (e.g. 16-mm pipe) become vulnerable, and the corresponding failure mode of the branch pipe depends on the restraint condition at the end of the branch pipe. However, i f the same trunk pipe is attached to relatively large diameter branch pipe (e.g. 42-mm pipe), the failure occurs at the tee, i f a socket outlet type tee is used. In analyzing similar branch pipe configuration, depending on the anchoring length and the location of the tee, the total resistance of a pipe configuration against relative movement in the direction of trunk pipe can assumed to be contributed by (a) the shaft (axial) frictional resistance of the trunk pipe, (b) the anchoring resistance of the tee and, (c) the anchoring resistance offered by the branch pipe during lateral movement. (Accordingly, the axial frictional resistance along the pipe could be determined as mentioned in Chapter 4.) For numerical formulations, the interaction arising from anchoring resistance of the tapping tee can be represented by a bilinear soil-spring, in which the maximum anchoring resistance (F t) and the corresponding displacement (yh) are proposed to be obtained from the guidelines specified for lateral movement of pipes and vertical anchor plates (e.g. A S C E , 1984). This anchoring force at the tee wi l l decide the amount of movement of the tee depending on the stiffness of the trunk pipe. With the movement of the tee, the anchoring resistance of the branch pipe (Fb) wi l l contribute to the total pullout resistance of the pipe system. In such situations, the 155 anchoring resistance of an unrestrained 16-mm branch pipe could be obtained from Equation 5.3 (subjected to an ultimate value of 2.5 kN) for known displacements of the tee. However, i f these deformations in the branch pipes are large enough to mobilize the restraint conditions at the branch pipe end, the analysis should now consider the types of restraining conditions at the end of the branch pipe and capacity of the branch pipe/ tee to undergo bending and tension. For example, in 42-mm branch pipes, when this restraint condition is mobil ized, the corresponding increase in resistance in the trunk pipe can be represented by Equation 5.4, subject to an ultimate value of 4.5 k N which corresponds to the capacity of the socket outlet tapping tee. These contributions from the branch pipe, together with contributions from other pipe components (e.g. tee and shaft friction in trunk pipe) would provide the necessary information to evaluate the response of a branch pipe configuration buried in dense sand. Further studies are required to estimate the anchoring resistance offered by the restrained small diameter (i.e. 16-mm) branch pipes and unrestrained large diameter (i.e. 42-mm) branch pipes. A comparison between the branch pipe tests performed with loose and dense sand backfills suggested that when branch pipes were not restrained from movement, the considerable increase in resistance was observed in dense sand than in loose sand. In contrast, the effect of soil density wi l l become less significant, i f the branch pipes were restrained from movement. 156 The results clearly highlighted that strain computations made using differential displacements at discrete locations on a pipe alone are not satisfactory and may lead to non-conservative assessment of piping vulnerability. The testing also provided valuable insight into technical detailing of pipe components (e.g. design of P V C protective sleeve and selection of tapping tees). In branch pipe configurations with large stiffness trunk pipe (i.e. 114-mm pipes) and low stiffness branch pipes (16-mm pipes), it was observed that the P V C protective sleeve tend to slip along the branch pipe with increasing ground displacements. Therefore, the sleeve would not provide the additional bending stiffness for the branch pipe. However, the capacity of the branch pipe could be increased by designing the P V C protective sleeve to properly lock onto the internal stiffener of the tapping tee during ground displacements. 6 . 3 Future research requirements From a utility owner's point of view, the findings from research work should eventually contribute to the solution of field problem. For example, the analysis approaches developed should enable the evaluation of the operational fitness of a given systems (or network) of distributional pipelines in an area prone to landslides. The findings from the current research work have revealed the complexities associated with the soil-pipe interaction in P E pipes subjected to relatively slow ground movements. In consideration of this, and on account of lack of studies in subject area, it is clear that there is still a wide spectrum of topics is yet to be explored. The research needs given below have been 157 identified with the objective of advancing the knowledge from the current work and developing input for pipeline design practice. As may be noted, the proposed future work wi l l stem from the already conducted investigations on axial, lateral and branch pipe systems. 6.3.1 Research requirements identified for axial soil loading It is judged that the derived closed-form solution in Section 4.1.2. has a string basis to form the framework to assess the field performance of P E pipes subjected to relative axial loading. The suitability of the present solution has been demonstrated based on some basic assumptions on bilinear pipe-soil interaction with average 5 (interface friction angle) and K (lateral earth pressure coefficient) values. With this intent, the closed-form solution needs to be further refined especially by recognizing the actual behavior of 5 and K with respect to increasing shear strain, dilation and pipe diametric reduction/expansion, and use this information in deriving a more representative interaction model in place of bilinear model presented herein. Although not included in the current derivation, the creep/stress relaxation behavior of P E would affect the response of the pipe system, especially when subject to slow intermittent ground displacements. Such aspects are needed to be incorporated in respective analytical formulations. 158 It is also proposed that work needs to be undertaken to develop charts directly usable by design/filed engineering personnel. For example pipe performance charts would yield information on the condition of the pipe (strain, force, etc) for ground displacements obtained from field surveys. For potential numerical analysis of complicated pipe configurations (e.g. modeling of a network of pipeline or a localized configuration), "equivalent" axial soil-spring characteristics are needed. The above refinements should allow the development of more representative soil-springs than what is used currently. It is proposed that, field pullout tests be conducted to verify some of the interaction models and the applicability of the closed-form solution for large relative displacements, (i.e. for large mobil ized frictional lengths of pipe). In addition, influence of soil aging could also be monitored from such tests. Unlike in rigid steel pipes, since gas pressure could have significant impact on the performance o f small diameter P E pipes, pullout tests are required to detect its impact due to (a) reduced diametric reduction, and (b) additional stiffness of the pipe in turn on the axial pullout resistance of the pipe. 6.3.2 Research requirements identified for lateral soil loading Another important component in understanding the pipe-soil interaction problem is the response of pipes subjected to ground movement perpendicular to the pipe axis (lateral 159 direction). It is important to note that, for a given pipe, the resistance created by the lateral movement is generally few magnitudes larger than the corresponding resistance created from axial movement. Currently, the available database for this aspect is very limited. Therefore, pullout tests are needed to investigate the impact of typical parameters that influence the pipe-soil interaction (e.g. impact of burial depth, soil density, soil dilation, etc). Again unique to these flexible pipes, factors such as gas pressure inside the pipe and pipe elongation would also influence performance, and the impact of such factors needs to be determined using full-scale tests. Indirectly, such experimental analysis would provide the means of ensuring the validity of using the already developed interaction models for steel pipes. 6.3.2 Research requirements identified for soil loading on piping networks. The fundamental behavioral characteristics derived from basic axial and lateral tests should be extended to develop solution methods for branch pipe systems. In general, the analysis of branch pipe systems is difficult due to complexities in interaction. In analyzing such systems, deriving analytical formulations from fundamental pipe-soil interactions sometimes would become difficult. A s a result, especially in ascertaining the interaction occurring at the pipe connection, fragility curves derived from experimental results would provide a reasonable means of analysis. It is recommended that, further research is needed to account the localized strains developed at the vicinity of the tapping tee connection. A lso in determining the total 160 pullout. resistance of a pipe system, the anchoring resistance of an unrestrained large diameter branch pipes (e.g. 42-mm pipe), and the anchoring resistance of restrained, small diameter branch pipe (e.g. 16-mm pipe) needs to be evaluated through appropriate full-scale model testing. It is also judged that appropriate 3-dimensional numerical analysis would provide a convenient means of analyzing such complex systems. 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(1996) "Recent Advances in the Modeling of Longitudinal Pipeline/Soil Interaction for Cohesive Soi ls" , Proceedings of the 15 t h International Conference on Offshore Mechanics and Arct ic Engineering, A S M E , Volume V , Pipeline Technology, pp. 127-139. 168 Rizkal la, M . , Poorooshasb, F. and Clark, J (1992) "Centrifuge Model ing of Lateral Pipeline/Soil Interaction", Proceedings of the 11 t h International Conference on Offshore Mechanics and Arct ic Engineering, A S M E , V o l . V , Pipeline Technology, Calgary, A B . Rowe, R.K. and Davis, E .H . (1982) "The Behavior of Anchor Plates in Sand", Geotechnique, V o l . 32, No . 1, pp. 25-41. Sivathayalan, S. (2000) "Fabric, Initial State and Stress Path Effects on Liquefaction Susceptibility of Sand", PhD Thesis, Department o f C i v i l Engineering, The University of British Columbia, Vancouver, B .C. Spangler, M . G . and Handy, R.L . (1982) "So i l Engineering" 4 t h edition, Harper & Row, New York. 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Proceedings of the 13 t h International Conference on Offshore Mechanics and Arctic Engineering, A S M E , N .Y . , pp. 127-139. Uthayakumar, M . (1996) "Liquefaction of Sands under Mult i-axial Loading", Ph.D Thesis, The University of Brit ish Columbia, Vancouver, B .C. Vaid. P. and Thomas, J . (1995) "Liquefaction and Post-liquefaction Behavior of Sand", Journal of Geotechnical Engineering, A S C E , Vo l . 121, No . 2, pp. 163-173. Varnes D. J . (1958) "Landslide Types and Processes", Editor: Landslide and Engineering Practice, Special Report 29, Washington, D C , Highway Research Board, National Research Counci l , pp. 20^17. Weerasekara, L., Wijewickreme, D. and Mitchel l , A . (2006) "Response of Tee-Junctions in Buried Polyethylene Natural Gas Distribution Piping Subject to Ground Movement" Proceedings of the 59 t h Canadian Geotechnical Conference, Vancouver, B .C . 170 Wijewickreme, D., Honegger, D.G. , Mitchel l , A . , and T.P. Fitzel l (2005) "Seismic Vulnerability Assessment and Retrofit of a Major Natural Gas Transmission System: A Case History", Earthquake Spectra, Journal of the Earthquake Engineering Research Institute (EERI), V o l . 21, No . 2, pp. 539-567. Wijewickreme, D., Karimian, A . , Honegger, D .G. (2005) "Effectiveness of Some Methods of Reducing Ax ia l Soi l Loads on Buried Pipelines Subjected to Ground Movements", Proceedings of 55 t h Canadian Geotechnical Conference, Saskatoon. Yang, Q.S and Poorooshasb, H.B. (1997) "Numerical Model ing of Seabed Ice Scour". Computers and Geotechnics, V o l . 21, No . 1, pp. 1-20. Zhou, J.Z. and Harvey, D.P. (1996) " A model for Dynamic Analysis of Buried and Partially Buried Piping Systems". Pressure Vessels and Piping Systems, P P V , A S M E , pp. 21-29. 171 APPENDIX A PIPE CONNECTIONS: SADDLE FUSION TAPPING TEES In gas distribution networks, pipe connections are common and they may provide vulnerable locations impacting the overall response of the pipe system. Typically, the branch pipe connections are made with equal and reduced outlet tees, wyes, or crosses, and two of the most common connections are: 1. Saddle fusion tapping tee type connection; 2. Butt tee type connection. The tapping tee connections are widely used in gas distribution industry due to ease of connection (see Figure 1-1). A lso shown are the two types of tapping tees that are commonly available, i.e. butt outlet type and socket outlet type. Typical ly, the tapping tees are installed with a " P V C protective sleeve" intended to protect the branch pipe from excessive shear and bending in the vicinity of the tapping tee outlet. Field saddle fusions are limited to 114 mm (4") diameter and smaller branch pipe connections on 300 mm (12") diameter or smaller mains. Complete details of such types of tapping tees and 172 available branch and trunk pipe connections could be found in respective manufacturer's catalogs (Performance Pipe, 2006). The pipe connections are generally intended to have higher strength capacity than the adjoining pipes. (c) Figure A . l (a) Tapping tee connection in an event of a ground movement, and cross section of a typical (b) 'socket outlet" tapping tee (b) "butt outlet" tapping Pipe failures in buried PE pipes In P E pipes, the failure mode of a non pressurized P E pipe can be of various forms, i.e. excessive pipe deflection, wall bucking, wall crushing, longitudinal bending, stress concentration, yielding. (Gazagnaire, 2000). However, depending on application and given loading condition, only few of these failure modes become important to investigate. Also deciding the ultimate failure of the plastic pipes would become difficult O'Rourke, 1990). The ultimate/bearing strain wi l l depend on small flows in the material and may see 173 a great deal of variation even under controlled lab conditions. Due to lack of field experience in failure modes especially in P E pipes has led to the question of defining suitable performance limit. Generally, a strain based criteria C S A (1992), is widely adopted for pipelines when determining the performance limits of a pipeline. In P E pipes, despite the better ductile behavior in short-term, many of the field failures are found to be attributed to brittle-like failures due to stress concentrations. Although, M D P E pipes show better stress cracking resistance than H D P E , brittle-like failures are still very common ( N T S B , 1998). These quasi-brittle cracks could occur even at a stress below the yield stress of the material. In field installations, the stress concentrations could arise due to external loading or due to presence joints or bends in the pipe system. The stress intensifications due to former wi l l become imperative in distribution networks, as joints are common in such pipe networks. Conversely, the stress cracking due to point loading can be minimized from proper installation practices. 174 APPENDIX B A X I A L A N D LATERAL PIPE PULLOUT TESTS RESULTS This appendix presents the test results of four axial pipe pullout tests and two lateral pipe pullout tests. The test configurations, load-displacement characteristics, and strain distributions and readings are also presented. 175 B . l . TestNo.60D SP1 SP2 SG1 SG2 / Load Cell 60 mm pipe Figure B. 1.1 Test configuration for Test No . 60D 176 5.00 0.00 100 200 300 400 SP1 Displacement, x (mm) 500 600 Figure B. l .2 Load-displacement characteristics for Test No . 60D 600 " i " 400 E. § 300 E a> 1 200 { in -100 100 Leading End (SP1) SP2 @ Trailing End 200 300 400 500 600 SP1 Displacement (mm) gure B . l .3 Displacement of the trailing and leading end of the pipe in Test No. 60D 177 Figure B.l .4 Radial and axial strain measurements at the leading end of the 60 mm pipe in Test No. 60D 178 B.2. Test No. 60D-1 .SP1 .SP2 .SP3 ,SP4 ,SP5 r Load Cell SG1 SG2 SG3 1.0 0.75 - x -0.75 T 60 mm pipe Figure B.2.1. Test configuration for Test No . 60D-1 179 0 - I— 1 —'— 1 — 1 —i— 1 — 1 — 1 — 1 —i— 1 — 1 — 1 — 1 —i— 1 — 1 — 1 — 1 —i— 1 — 1 — 1 — 1 —i— 1 — '— 1 —'—i 0 100 200 300 400 500 600 SP1 Displacement, x (mm) Figure B.2.2. Load-displacement characteristics for Test No. 60D-1 600 SP1 Displacement (mm) Figure B.2.3. Displacement of several locations of pipe in Test No. 60D-1 180 • Radial Strain Axial Strain 400 500 600 SP1 Displacement (mm) Figure B.2.4 Radial and axial strain measurements at the leading end of the 60 mm pipe in Test No . 60D-1 Figure B.2.5. Strain distribution along the 60 mm pipe for different displacements of the leading end 181 B.3. TestNo.ll4D SP1 SP2 SG1 SG2 / Load Cell 114 mm pipe Figure B.3.1. Test configuration for Test No . 114D 182 8 Figure B.3.3. Displacement of the leading and the trailing ends of the pipe in Test No . 114D 183 6000 T 5000 4000 3000 1 2000 -c 1 -4—' 1000 -CO 0 -Radial Strain Axial Strain 300 400 500 600 SP1 Displacement (mm) Figure B.3.4. Radial and axial strain measurements at the leading end of the 60 mm pipe in Test No. 114D SP1 Displacement, x (mm) Figure B.3.5. Load -displacement characteristics for the second loading (pushing direction) in Test No. 114D 184 300 — Radial Strain ('+'; — Axial Strain ('-') 400 500 600 -8000 SP1 Displacement (mm) Figure B.3.6. Radial and axial strain measurements at the leading end of the 114 mm pipe in second loading (pushing direction) Figure B.3.7. Stress-strain diagram derived from strain gauge and load cell readings in axial pullout tests 185 B.4. Test No. 114D-1 ,SP1 .SP2 1 u 1 Load Cell SG1 SG2 1.0 ^ ^ 0.75 SG3 T - X -0.75 114 mm pipe Figure B.4.1. Test configuration for Test No . 114D-1 186 o u c S w "5> o: 3 o 3 CL 100 200 300 400 SP1 Displacement, x (mm) 500 600 Figure B.4.2. Load-displacement characteristics of Test No . 114D 600 500 E 400 E. § 300 E tu | 200 w % 100 -100 Leading End (SP,,;Y SP2 @ Trailing End 100 200 300 400 500 600 700 SP1 Displacement (mm) gure B.4.3. Displacement of the leading and the trailing ends of the pipe in Test No. 114D-1 187 8000 6000 4 -4000 -1 — - — - - • - J SP1 Displacement (mm) Figure B.4.4. Radial and axial strain measurements at the leading end of the 114 mm pipe in Test No. 114D-1 6000 -, Length along the pipe (m) Figure B.4.5. Strain distribution along the 114 mm pipe for different displacements of the leading end 188 B.5. Test No.l 14D (concrete) SP1 SP2 SG1 SG2 Load Cell 114 mm pipe filled with concrete Figure B.5.1. Test configuration for Test No . 114D-1 189 12 2 -' 0 I ' 1 1 1 I 1 1 1 1 i 1 1 1 ' I i . j i i . i 0 100 200 300 400 500 600 SP1 Displacement, x (mm) Figure B.5.2. Load-displacement characteristics of Test No . 114D (concrete) gure B.5.3. Displacement of the leading and the trailing ends of the pipe in Test No. 114D (concrete) 190 2500 2000 4 -1000 -J- - - - — J SP1 Displacement (mm) Figure B.5.4. Radial and axial strain measurements at the leading end of the 60 mm pipe in Test No . 114D (concrete) 191 B.6. Test No.l 14D- (2.5m) Figure B.6.1. Test configuration for Test No . 114D- (2.5m) 192 14 12 Z 10 CD O c S in 8 4 in o Q£ 6 +J 3 O 2 4 2 4 100 200 300 400 SP1 Displacement, x (mm) 500 Figure B.6.2. Load-displacement characteristics of Test No . 114D- (2.5m) 600 -100 Length along the branch pipe (mm) Figure B.6.3. Displacement along the pipe in Test No . 114D-(2.5 m) test 193 0 -10 10 mm - * - 5 0 mm —(—100 mm 200 mm 300 mm - * - 4 0 0 mm 500mm ffi 200 400 600 800 1000 1200 1400 1600 1800 2000 2200 Length along the branch pipe (mm) gure B.6.4. Bending moment for different displacements of the pipe in Test No . 114D-(2.5m) 194 B.7. Test No.l 14D-(1.5m) SP1 111 L n 1.125 SP2 easurement I - S G m points T Z T Load Cells r e Figure B.7.1. Test configuration for Test No . 114D-(1.5m) 195 a> u c a v> Jo <D OH *J 3 o 3 Q. 100 200 300 400 SP1 Displacement, x (mm) 500 Figure B.7.2. Load-displacement characteristics of Test No . 114D-(1.5m) 20 X 1 * — * • 200 400 600 800 1000 1200 Length along the branch pipe (mm) 1400 Figure B.7.3. Bending moment for different displacements of the pipe in Test No. 114D-(1.5m) 196 APPENDIX C B R A N C H PIPE PULLOUT TESTS RESULTS This appendix presents the test results and configurations of eight branch pipe pullout tests. The test results of two pullout tests on 114 and 60-mm pipes with tapping tees are also presented. Eventhough, the tests results of the butt-tee connection is not discussed in this thesis, results are presented herein. 197 C.l. Test No. 60D16-U SP1 P ft Load Cell SP2 60 mm Trunk PiDe SP3 i i i iyrr SG1 SG 1.0 0.75 >< > 9. i r r r 16 mm branch Pipe SG3 SG4 SG5 0.75 0.75 < >< > 3 u • SG6 \ String Potentiometers SP4 Figure C. 1.1 Test configuration for Test No . 60D16-U 198 Figure C . l . 2 . Load-displacement characteristics of Test No . 60D16-U 600 SP1 Displacement (mm) Figure C . l . 3 . Displacement at different locations of the trunk pipe 199 80000 -20000 x - 1 Length along the pipe (mm) Figure C. 1.4 Strain distribution along the trunk pipe for different displacements of the leading end Figure C. 1.5 Distribution along different locations along the trunk pipe for different displacements of the leading end 200 C.2. Test No. 60D16-R 60 mm Trunk Pine SP1 SP2 SP3 [f lO: Load Cell 8 SG1 SG 1.0 0.75 >< > I T SG3 SG4 SG5 0.75 0.75 < >< > n SG6 16 mm branch Pipe Load Cells Figure C.2.1 Test configuration for Test No . 60D16-R 201 100 200 300 400 SP1 Displacement, x (mm) 500 Figure C.2.2 Load-displacement characteristics of Test No . 60D16-R 600 500 4-E E 400 c CD E 300 o o CO -100 Leading End (SP1K 100 200 300 400 500 600 SP1 Displacement (mm) Figure C.2.3 Displacement at different locations of the trunk pipe 202 gure C.2.4 Distribution along different locations along the trunk pipe for different displacements of the leading end 203 C.3 Test No.l 14D16-U SP1 SP2 114 mm Trunk Pipe SP3 Load Cell SG1 SG 1.0 0.75 X > 9, l l l ly J I I II 16 mm branch Pipe SG3 SG4 SG5 0.75 0.75 < X > LT u 0 SG6 SP4 String Potentiometers Figure C.3.1 Test configuration for Test No . 114D16-U 204 100 200 300 400 SP1 Displacement, x (mm) 500 600 Figure C.3.2 Load-displacement characteristics of Test No . 114D16-U 700 SP1 @ First Tee S P 4 i Trailing End S P 2 @ econd Tee 1st Branch Pipe: Trailing End 2nd Branch Pipe: Trailing End 600 700 SP1 Displacement (mm) Figure C.3.3 Displacement at different locations of the trunk and branch pipe 205 20000 -10000 Length along the pipe (m) Figure C.3.4 Strain distribution along the trunk pipe for different displacements of the leading end 0.8 <u 0.6 o La £ I 0.4 0.2 — At second tee — At first tee 100 150 200 250 SP1 Displacement (mm) 300 350 400 Figure C.3.5 Tensile forces calculated from strain gauge readings at a distance of 25 mm from the internal stiffener of the first and the second tee 206 C A 114D16-RTEST SP1 q U Load Cell SP2 114 mm Trunk PiDe SP3 SG1 SG; 1.0 0.75 —><—> I 16 mm branch Pipe SG3 SG4 SG5 0.75 0.75 49* SG6 SP4 Load Cells Figure C.4.1 Test configuration for Test No . 114D16-R 207 20 0 I i | I I I I | 1 I 11 I I "IL-p 0 100 200 300 400 500 SP1 Displacement, x (mm) Figure C.4.2 Load-displacement characteristics of Test N o . 114D16-R Figure C.4.3 Tensile force at the ends of the branch pipe and close to the internal stiffener of the tee 208 700 600 _ 500 E E « 400 c 9 | 300 -100 100 Leading End (SP1) 200 300 400 SP2 @ First Tee SP3 @ Second Tee 500 600 700 SP1 Displacement (mm) Figure C.4.4 Displacement at different locations of the trunk pipe 20000 15000 10000 "ST c "«5 5000 -5000 -10000 500 1000 1500 N 2000 First Tee Edge of the box 2500 3 0 5 ° \ 3 5 0 0 Second Tee Edqe of the box Length along the pipe (m) Figure C.4.5 Strain distribution along the trunk pipe for different displacements of the leading end 209 C.6 TestNo.ll4L42-U SP1 n r 7 Load cell SG1 SP2 < — | ^ SG3 1.0 SG4 1.0 1.0 42 mm Branch Pipe string Dotentiometer A SG5 & 6 0.5 ft «Y SG6 & 7 SP3 114 mm pipe Figure C.5.1 Test configuration for Test No . 114L42-U 210 10.00 o o c 2 <2 '35 0) CH +J 3 o 3 0. 9.00 8.00 7.00 6.00 5.00 4.00 3.00 2.00 1.00 0.00 0.00 -J—I—' 100.00 200.00 300.00 400.00 SP1 Displacement, x (mm) 500.00 Figure C.5.2 Load-displacement characteristics of Test No . 114L42-U 700 -100 700 SP1 Displacement (mm) Figure C.5.3 Displacement at different locations of the trunk and branch pipe 211 8000 -6000 Length along the pipe (m) Figure C.5.5 Bending moment calculated from strain gauge readings in the branch pipes 212 Figure C.5.6 Tensile forces calculated from strain gauge readings in the branch pipes 213 C.6Test N0.114L42-R 1.0 SG1 SP2 4-—I _ SG3 =LTJp£rz: 1.0 SG4 1.0 JT- SG5 & 6 0.5 0 «T SG6 & 7 42 mm Branch Pipe Load cell Figure C.6.1 Test configuration for Test No . 114L42-R 214 0 100 200 300 400 500 SP1 Displacement, x (mm) Figure C.6.2 Load-displacement characteristics of Test N o . 114L42-R 600 -100 J SP1 Displacement (mm) Figure C.6.3 Displacement at different locations of the trunk pipe 215 25000 -5000 1 Length along the pipe (m) Figure C.6.5 Bending moment calculated from strain gauge readings in the branch pipes 216 5 0 100 200 300 400 500 600 SP1 Displacement (mm) Figure C.6.6 Tensile forces calculated from strain gauge readings in the branch pipes 217 C.7 TestNo.ll4D42-U SP1 7 ZLTZ SG1 SP2 4. • 4 — 1 ^ SG3 1.0 SP3 SG4 ZLTZ Load cell 1.0 1.0 A SG5 & 6 0.5 rji «T SG6 & 7 42 mm Branch Pipe 114 mm pipe string Dotentiometer Figure C.7.1 Test configuration for Test No . 114D42-U 218 100 200 300 400 SP1 Displacement, x (mm) 500 Figure C.7.2 Load-displacement characteristics of Test No . 114D42-U E E +J c <D E o re Q. (0 Q a -100 SP3 @ Trailing end SP4 @ branch pipe end 100 200 300 400 500 600 700 SP1 Displacement (mm) Figure C.7.3 Displacement at different locations of the trunk pipe 219 8000 7000 6000 5000 1 4000 c 3000 ' r c i_ <f> 2000 1000 0 -1000 -2000 Leading end Trailing end_ - •—2 mm - * - 5 mm —"—10 mm ——20 mm 50 mm Length along the pipe (m) Figure C.7.5 Bending moment calculated from strain gauge readings in the branch pipes 220 221 C.8TestNo. 114D42-R 1.0 SG1 SP2 — | ~ _ SG3 ==nQn= 1.0 SG4 1.0 yC~ SG5 & 6 0.5 r}j < Y SG6&7 42 mm Branch Pipe Load cell Figure C.8.1 Test configuration for Test No . 114D42-R 222 18 16 z 1 4 ^ 12 o c 2 10 .2 M a: 8 3 o "5 DL 6 4 4 Load cell reading at the leading end Load cell reading at the end of the branch pipe 100 200 300 400 SP1 Displacement, x (mm) 500 Figure C.8.2 Load-displacement characteristics of Test N o . 114D42-R -100 SP3 @ Trailing end 100 200 300 400 500 600 700 SP1 Displacement (mm) Figure C.8.3 Displacement at different locations of the trunk pipe 223 10000 7 -2000 J - — -Length along the pipe (m) Figure C.8.4 Strain distribution along the trunk pipe for different displacements of the leading end Figure C.8.5 Bending moment calculated from strain gauge readings in the branch pipes 224 Figure C.8.6 Tensile force calculated from strain gauge readings and load cell measurement at the end of the branch pipe. 225 C.9 Test No. 60D16 (tee only) Load Cell 60 mm Trunk PI'DG SP1 SP2 SP3 n a I LT SG1 SGZ'-'SGS SG4 SG5 •—' SG6 1.0 0.75 0.75 0.75 X > < X > Figure C.9.1 Test configuration for Test No . 60D16 (tee only) 226 12 10 cu o c CD 4-1 CO '55 6 CD a. *-> o 1 4 0. 100 200 300 400 SP1 Displacement, x (mm) 500 600 Figure C.9.2 Load-displacement characteristics of Test No. 60D16 (tee only) o -100 SP2 @ First Tee SP3 ©Second Tee Trailing End (SP4) 100 200 300 400 500 600 700 SP1 Displacement (mm) Figure C.9.3 Displacement at different locations of the trunk pipe 227 60000 50000 40000 30000 .E 20000 2 10000 0 -10000 -20000 30(|0\ 3500 Second Tee Leading end Trailing end Length along the pipe (m) Figure C.9.4 Strain distribution along the trunk pipe for different displacements of the leading end Figure C.9.5 Distribution along different locations along the trunk pipe for different displacements of the leading end 228 C.10 Test No. 114D42 (tee only) Load Cell 114 mm Trunk PiDe SP1 SP2 SP3 TL TL SG1 S G ^ S G S SG4 S G 5 u SG6 1.0 0.75 0.75 0.75 ><—> < X > Figure C. 10.1 Test configuration for Test No . 114D142 (tee only) 229 100 200 300 400 S P 1 Displacement, x (mm) 500 Figure C.10.2 Load-displacement characteristics of Test No . 114D42 (tee only) E £ .*-» c o E CD U re a. cn b a. 700 600 500 400 4 300 200 100 0 -100 SP2 @ Tee Leading End (SP1) SP3 @ Trailing End 100 200 300 400 500 600 700 SP1 Displacement (mm) Figure C.10.3 Displacement at different locations of the trunk pipe 230 3 / > « Tapping Tee 3000 Trailing end 3500 • Length along the pipe (m) Figure C . l 0.4 Strain distribution along the trunk pipe for different displacements of the leading end 2 3 1 C. 11 Test No. 114D42 - U (butt tee) SP2 <_ 1.0 S G 3 I E Z S G 4 SG1 S G 2 1.0 1.0 — SG5 & 6 0.5 «Y S G 6 & 7 42 mm Branch Pipe string D o t e n t i o m e t e r Figure C. 11.1 Test configuration for Test No . 114D42-U (butt tee) 232 100 200 300 400 SP1 Displacement, x (mm) 500 Figure C . l 1.2 Load-displacement characteristics of Test No . 114D42-U (butt tee) 700 600 + 700 SP1 Displacement (mm) Figure C . l 1.3 Displacement at different locations of the trunk pipe 233 .S 5000 co 35 o -5000 -10000 -15000 -20000 500 1000 1500 / 20] Butt-tee Leading end 2 mm -*- 5 mm 10 mm — 20 mm — 50 mm 100 mm — 200 mm — 300 mm 400 mm 3000 3500 Trailing end. Length along the pipe (m) gure C. 11.4 Strain distribution along the trunk pipe for different displacements of the leading end 234 APPENDIX D PHOTOGRAPHS This appendix presents the photographs of test preparation (couplings, strain gauging, etc), test setup and pipe failure/ deformation modes (after testing) for various pipe configurations. 235 Figure D.3 Strain gauges secured to the pipe with plastic ties. Figure D.4 Unrestrained 42 mm pipe (in Test No. 114L42-U) with string potentiometer wire attached to the pipe end. 237 Figure D.5 Restrained 42 mm pipe (in Test No. 114L42-R) with a load cell attached to the pipe end. Figure D.6 After placing the pipe inside the soil box in Test No . 60D16-U. 238 Figure D.7 Deformation of the branch pipe in Test No . 60D16-U. Figure D.8 Deformation (buckling) of the branch pipe closer to the tee in Test No . 60D16-U 239 Figure D.9 Yielded branch pipe in Test No . 114D16-U Figure D . l 1 Deformed branch pipe after testing in Test No . 114L42-U Figure D. l 3 Signs of yielding underneath the tee in Test No . 114L42-R Figure D. 14 Formation of a plastic crack in Test No . 114D42-U 242 Figure D.15 Separated 42-mm branch pipe after testing in Test No . 114D42-R 

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