RESPONSE OF TEE-JUNCTIONS IN BURIED...

RESPONSE OF TEE-JUNCTIONS IN BURIED POLYETHYLENE NATURAL GAS DISTRIBUTION PIPING SUBJECT TO GROUND MOVEMENT Lalinda Weerasekara, Dept. of Civil Engineering, University of British Columbia, Vancouver, BC, Canada Dharma Wijewickreme, Dept. of Civil Engineering, University of British Columbia, Vancouver, BC, Canada Allen Mitchell, Terasen Gas Inc., Surrey, BC, Canada ABSTRACT The performance of buried natural gas pipeline systems in areas subjected to ground deformations is an important engineering consideration for the utility owners. The response of tee-junctions in distribution systems with polyethylene (PE) piping (typically comprising trunk pipe, “tapping tee”, and branch pipe) under ground movements is complex and one of the areas of particular interest. Full-scale physical modeling of buried PE pipe tee-junctions under pullout loadings were undertaken to investigate this soil-pipe interaction problem. Test results demonstrate that the physical location of vulnerability would depend on relative pipe sizes, soil density, and boundary conditions. In essence, depending on the configuration, large pipeline strains or failure would be localized on branch pipe or trunk pipe, or at the “tapping tee” connection. The test results directly contribute to developing a reliable database for developing/validating numerical models to capture the response of pipeline distribution systems subjected to ground movements. RÉSUMÉ L’exécution de systèmes enterrés de canalisation de gaz naturel dans les secteurs exposés aux déformations du sol est une considération de technologie importante pour les propriétaires du service. La réponse des pièces en tee junction dans des systèmes de distribution avec des tuyaux en polyéthylène (PE) (comportant typiquement le tuyau de tronc, le ‘tapping tee’, et le tuyau de branche) soumis à des mouvement de sol est complexe et l’un des secteurs de particulier intérêt à cet égard. Modèles physiques de tuyaux PE à pleine grandeur soumis à des chargements d’arrachement ont été entrepris afin d’étudier ce problème d’interaction sol-tuyau. Les résultats d’essais démontrent que l’emplacement physique de la vulnérabilité dépendrait de la taille relative des tuyaux, la densité du sol, et les conditions de frontière. Essentiellement, selon la configuration, grandes déformations ou rupture du tuyau serait localisées sur le tuyau de branche ou le tuyau de tronc ou dans la connexion ‘tapping tee’. Les résultats d’essai contribuent directement à développer une base de données fiable pour développer/valider models numériques pour capturer la réponse de systèmes de distribution soumis aux mouvements de sol. 1. INTRODUCTION

Since its introduction in the late 1960’s, polyethylene (PE) pipes are becoming increasingly popular in natural gas distribution networks. In North America, 90% of the natural gas distribution systems comprise plastic pipes, of which, 99% are PE pipes (PIPA, 2001). PE pipes have become more popular due to its lower material, installation and maintenance costs, corrosion resistance, lower friction at the interface, lightweight, and its claimed greater capacity to accommodate displacements than its counterpart, steel pipes. Although HDPE (High Density Polyethylene) pipes are also used in practice; MDPE (Medium Density Polyethylene) pipes account for 2/3‘s of the usage in gas distribution industry. MDPE pipes have the added advantage of having higher flexibility and fracture toughness while having comparable long term strength and stiffness to that of HDPE (Stewart et al, 1999). The majority of PE pipe failures in the field are attributed to point loading and mechanical damages (PIPA, 2001). Nevertheless, damage to piping systems due to stress and strain concentrations induced from ground movements is also a major concern to pipeline owners. In the evaluation of the performance of PE pipes in the presence of ground movements, current practice essentially relies on the analysis methods developed for rigid steel pipes. Considering the relatively smaller deformation stiffness

and increased potential for time-dependent response (creep) of PE as a pipe material, there is a good likelihood for significant limitations in the use of methods developed for steel pipes in the assessment of the response of PE pipes. Over the years, numerous studies have been performed on steel pipes to characterize the behavior when the pipes are subjected to ground movement. These involve, laboratory tests spanning from centrifuge tests to full-scale model tests, field testing and monitoring, numerical and analytical modeling. Despite the advances in steel pipe design, there have been very few studies carried out to characterize the performance of the PE pipe. With this background, a detailed research program involving full-scale physical modeling has been undertaken at the University of British Columbia (UBC), Canada, in partnership with Terasen Gas Inc., B.C., Canada, to study soil-pipeline interaction during permanent ground displacements affecting distribution networks of PE piping. The objectives of this research are to determine the contributing factors, and key parameters influencing the behavior of polyethylene pipe buried in soil, and provide input to developing a suitable set of guidelines and criteria to increase the level of confidence in evaluating the operational fitness of a given PE piping system.

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(a)

(b)

Tapping Tee

(c)

Saddle fusion joint

Trunk Pipe

Branch Pipe

PVC Protective Sleeve

PVC Protective Sleeve

Figure 1: (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 tee.

X

Y

Previously, Anderson et al (2004) examined the behavior of straight pipes and branch PE pipes when ground movement is occurring along the axis of the trunk line. The full-scale tests performed included, axial pullout tests of straight pipes of diameters (D) 114 and 60 mm, and some limited pullout tests on branch pipe. Anderson, et al. (2004) observed that in branch pipe tests, the soil load developed due to the branch or “branch contribution” (i.e., difference in pullout resistance between straight pipe and the branch pipe) was independent of the trunk pipe diameter, for the same size of branch pipes. In addition, pullout tests have been performed at Cornell University (Stewart et al, 1999) to explore the effects of temperature changes and effect of temperature induced cyclic loading on PE pipes. The pullout resistance of the PE pipes was measured in a temperature controlled test compartment at different temperatures. Results from these tests have indicated that temperature changes can cause significant change in pipe diameter, in turn, affecting the soil friction on the pipe. During permanent ground movement, the performance of a pipe system 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. Furthermore, the soil-pipe interaction occurring in a branch pipe system especially close to the joint can be very complex (Anderson, 2004). In the present study, considering the need to obtain a more thorough understanding of the performance at the tee-junctions, detailed physical model tests were undertaken on buried

“tapping tee” configurations of PE piping (i.e., tapping tees join the branch pipes to the trunk line with a saddle fusion joint as shown in Figure 1). Typically, the tapping tee is also provided with a “PVC protective sleeve” to reinforce the branch pipe from shear forces and bending at the vicinity of the tapping tee outlet [see Figure 1(b) and 1(c)]. A full-scale testing program was designed to account for different pipe diameters, pipe configurations and soil densities (loose and dense), and this paper describes the experimental procedure and results. 2. TEST SETUP & EXPERIMENTAL DETAILS The test setup mainly comprise a 3.8m L x 2.5m W x 2.5m H soil box, a hydraulic actuator system having a pulling load capacity of 220 kN, and a data acquisition system. The length of the soil box was selected to develop sufficient strain along the length of the trunk pipe. In addition, the width of the soil box was such that it could accommodate a reasonable length of branch pipe to study its deformations and anchoring. For all tests, dry Fraser River sand (FRS), which has been extensively used in laboratory research at UBC over the past 10 years, was used as the soil backfill surrounding the pipe. The mineral composition of the FRS was found to be 40% quartz, 11% feldspar, 45% unaltered rock fragments and 4% other minerals. The minimum and maximum void ratios were 0.62 and 0.94 (Wijewickreme et al., 2005). The grain size distribution shows minimum and average grain sizes of 0.074 mm and 0.23 mm respectively. The coefficient of


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uniformity was found to be 1.5 (Anderson, 2004). Density measurements showed that the average density is 1610 kN/m3 in dense sand and 1520 kN/m3 in loose sand. For all experiments, MDPE pipes having a density of 0.940 g/cm3 have been used. Details of the pipes used in the tests are given in Table 1. Table 1. Details of PE pipes used in testing

Pipe/Tube diameter D, (mm)

Wall thickness (mm)

M.O.P* S.D.R**

114 (4”) 10.32 100 11 60 (2”) 5.48 100 11 42 (1 1/4”) 4.22 110 10 16 (1/2” tubing) 2.28 n.a n.a

* M.O.P: Maximum Operating Pressure ** S.D.R: Standard Dimension Ratio (Minimum outside diameter/Minimum wall thickness) The details of the tested pipe configurations are schematically shown in Figure 2. Strain gauges were mounted at selected locations, as indicated by the symbol “SG” shown in the figure. After mounting the strain gauges, the pipe was placed inside the sand box and connected to the hydraulic actuator via appropriate couplings. A load cell, attached in series, measured the pullout resistance of the pipe system. String potentiometers were also attached to selected locations of the pipe, e.g. at tapping tees, trailing end of the trunk and branch pipes etc. The sand backfill material was moved into the soil box using large storage bags (1 cu.m.). An overhead crane was employed to lift and position these sand bags at the desired location(s) in the chamber; it was possible to release sand through an opening at the bottom of the bags and, then, spread inside the chamber to form a layer by controlling the movement of the crane as desired. In preparing for tests with dense sand, each sand layer (of thickness 100-150mm) was leveled and compacted using four passes of a ½ tonne static roller each in lateral and longitudinal directions. The sand slumped from the bottom chute of the bag was used, without compaction, as the backfill to achieve a relatively low soil density for those tests with medium loose backfill condition. All the tests were conducted with burial depth to the spring line level of trunk line pipe (H) kept at 0.6 m; these depths are generally in line with the typical burial depths used in natural gas distribution systems. Thus, the H/D ratios for the test configurations having trunk pipe diameters of 114 mm and 60 mm were 5.2 and 10, respectively. The tests were conducted by pulling the buried trunk pipe from one end (referred to as the “leading end” herein) using the hydraulic actuator system while recording the pullout resistance, axial strain(s), and displacement(s) with the progress of pipe movement. Comprehensive details on the test setup, strain gauge selection and mounting techniques are given by Anderson (2004), and, therefore, they are not repeated herein. The pipeline configurations were subjected to pullout using a constant rate of displacement of 36 mm/hr applied at the leading end of the pipe. Study of the effect of

displacement rate on pullout resistance was beyond the scope of this study. Eight branch pipe tests were performed as listed in Table 2. 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 trunk pipe diameter 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 at locations market with an ‘A’ in Figure 2; whereas, “R” designates branch pipes were restrained against movement at location A, e.g. 60D16-U refers to 60mm diameter trunk pipe with unrestrained 16 mm branch pipes in dense soil.

In tests with unrestrained branch pipe, string potentiometers were used to record the displacement at the location A, whereas in the case of restrained pipe, load cells were used to restrain the trailing end of the branch pipes. Both unrestrained and restrained conditions were considered important to be tested. For example, in real-life pipe installations, small diameter branch pipes are generally laid with some amount of slack, allowing some axial movement in the pipe (Stewart, 1999); the unrestrained conditions discussed above can be considered similar to this field

Load Cell

Figure 2. Plan views of pipe configurations for (a) 60D16-* and 114D15-* tests, (b) 114*42-* tests; (SG = strain gauge). Note: All lengths are in meters.

(b)

A

A A

SG2 SG1 SG3 SG4

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Load Cells or string potentiometers

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Branch Pipe

SG6 SG5 SG4 SG3 SG2 SG1

0.75 0.75 0.75 0.75


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configuration. On the other hand, the branch pipes can be significantly restrained by connections to buildings, by pipe components (joints, bends etc) or by the longitudinal frictional anchorage of the pipe, thus demonstrating the need to test with the location A restrained. All tests were performed under general ambient laboratory temperature conditions, which is about 20 ± 2 °C. 3. RESULTS & DISCUSSION The experimental observations provide an opportunity to examine possible modes of failure prevailing under different boundary and soil/pipe conditions. The failure modes of a non pressurized PE pipe can be in various forms, e.g. excessive pipe deflection, wall bucking, wall crushing, longitudinal bending, stress concentration and yielding. (Gazagnaire, 2000). Depending on application and given loading condition, only some of the failure modes become important and need to be investigated. While the strain gauge readings allowed to identify the critical sections of the pipeline, since the reliability of axial strain gauge readings were limited to about 3-5%, it was not possible to obtain strains at the instance of rupture etc. In following section, the results will be discussed in relation to the three pipe configurations tested: (a) 114*42-* tests (b) 114D16-* tests and (c) 60D16-* tests.

3.1 Tests No. 114*42-* : 114mm trunk pipes with 42mm branch pipes

The test configuration is shown in Figure 2(a). The load (at the leading end) vs. displacement characteristics of two branch pipe systems (restrained and unrestrained) in loose sand are shown in Figure 3. In addition, the load measurement at the restrained end of the branch pipe is also shown.

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SP1 Displacement, x (mm)

Pu

llo

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)

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Load cell reading at end of the branch pipe

Unrestrained

Figure 3. Variation of pullout resistance of the (a) 114L42-U and 114L42-R tests (together with the load cell reading at the end of the branch pipe, in restrained test) In all tests, at small displacements (< 16 to 20mm), the strain readings on the trunk pipe showed strain distributions similar

to those observed by Anderson (2004) during the pulling of a straight pipe, i.e. linearly decreasing strain along the length of the pipe (see Figure 4). However, as the axial displacement of the leading end exceeded ~20mm, the axial strain in the trunk pipe at a location in front of the tee increases relative to the leading end of the pipe (NOTE: herein, “in front” from a certain point means on the “leading end” side from that point – vice versa - the word “behind” is used to denote locations on the side opposite to the “leading end”). In all cases, the strain gauges were attached to the crown of the trunk pipe. In a few tests, additional gauges were mounted to the pipe’s invert to examine potential variations in axial strain distribution at a given cross section. Identical strains were observed at the crown and invert levels when the gauges were located about 1 m away from the tee. However, immediately in front of the tee, larger strains were measured at the crown compared to the invert, signifying large stress concentrations, closer to the tee. Immediately behind the tee, the measured strains were small or may become even negative (compressive strains) at large displacements. Clearly, the anchoring effect due to the presence of the tee seems to provide a major discontinuity to strain development on the trunk pipe during the load transfer process to the remaining part of the trunk pipe beyond the tee.

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Figure 4. The strain distribution along the trunk pipe for different displacements of the leading end in 114D16-R test. The geometric discontinuation at the tee led to stress build-up and resulted in continual yielding and crack formation at the tee (see Figure 5). For all 114*42-* type tests, “socket outlet” tapping tees were employed. Therefore, in this type of tapping tee, the point where the cross section is smaller than the diameter of the branch pipe [see Point X, Figure 1(b)] was consistently noted to be the vulnerable location in tests conducted with this configuration. For example, in Test No. 114D42-R, complete separation of the branch pipe was observed at the tee, when the pipe was examined after exhuming. Cracking was also observed [figure 5(b)] in other tests to varying degree. As expected, in all restrained tests,

Trailing end

Leading end


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the crack was wider than in unrestrained test. Also the damage was greater in dense tests than in loose tests.

(a)

(b) Figure 5: (a) The plastic deformation of the tee in horizontal plane (b) Formation of a crack at the “socket type” tapping tee, in 114D42-U test (photos taken on exhumed pipe after testing). In tapping tee connections, the axis of branch pipe is connected with a vertical offset to the axis of the trunk pipe [Figure 1(b)]. Owing to this offset, especially when the branch pipe is restrained, the tapping tee connection can induce moment/rotation in the trunk pipe. This rotational mode in combination with the plastic deformations at the tee as discussed above (where the cross section of the tee is small) seems to have provided increased flexibility to the tee as shown in Figure 5 (b). This flexibility automatically available in the “socket outlet” type tapping tee seems to

have reduced the relative anchoring effect of the branch pipe, in turn, minimizing transfer of any significant stresses to the trunk pipe, in spite of the relatively large diameter and stiffness of the branch pipe. In this testing, the maximum strain measured in the trunk pipe was about 2%. However, it must be noted that this protection of the trunk pipe is achieved at the expense of the structural failure of the weak cross section at the tee 3.2 Tests No. 114D16-* : 114mm trunk pipes with 16mm branch pipes This test configuration is shown in Figure 2(b). The load (at the leading end) displacement characteristics of 114D16-* tests are shown in Figure 6. “Butt type” tapping tees were used to connect the trunk pipe to the 16mm branch pipe. The loads recorded at the ends of the branch pipes, in restrained tests are not shown for clarity. In Test No. 114D16-R, the observed sharp drop in load (at about 480 mm displacement) was due to separation of the branch pipe from the tee after excessive yielding. The strain gauge readings and string potentiometers attached to the branch pipes showed significant amount of deformation, confined to a region within 0.3 m from the trunk pipe axis. Very significant strains were recorded in branch pipes (closer to the tee) even after a slight amount of displacement in trunk pipe, signifying that 16-mm diameter branch pipes are greatly vulnerable even at small relative ground movements.

60D16-R

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Displacement of the leading end : SP1/(mm)

114D16-R

Pullout re

sis

tance/(

kN

) 114D16-U

60D16-U

Figure 6. Pullout resistance of the 114D16-* and 60D16-* tests along with the leading end displacement. In the case of 16-mm branch pipes, the observed modes of failure in unrestrained and restrained tests were different. When restrained, the 16-mm branch pipes showed significant amount of yielding before eventually separating at the tee. Inspection of pipes after exhuming, in unrestrained case, branch pipes showed buckling at Point Y [Figure 1(c)], thereby possibly reducing the pipe cross section available for gas flow.). It is further observed that the strains in the 114-mm trunk pipe were in the same order of magnitude as in


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114D42-* case. This level of strain is not sufficient to cause major concerns with regard to potential failure of trunk pipe. 3.3 Tests No. 60D16-* : 60mm trunk pipes with 16mm branch pipes Except for the difference in diameter of the trunk pipe, the test configuration and the type of tapping tees used in this test series was similar to that of 114D16-* [Figure 2(b)]. In Tests No. 60D16-U and 60D16-R, there was no displacement or load recorded at the ends of the branch pipes during the pullout of trunk pipe. In turn, the load cell at the leading end displayed almost identical load-displacement characteristics (Figure 6) .

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Leading end Trailing end

Second endFirst tee

Figure 7. The strain distribution along the trunk pipe for different displacements of the leading end in 60D16-U test.

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SP1 Displacement (mm)

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Dis

pla

cem

ent (m

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Leading End (SP1)

First tee

Second teeTrailing end

Figure 8. Displacement of tees and trailing end with respect to the displacement of the leading end (in 60D16-U test.)

The strain distributions obtained from strain gauge readings in Test No. 60D16-U are plotted in Figure 7. Although not plotted for clarity, the strains observed in Test No. 60D16-R were also very similar to that given in Figure 7. These observations also indirectly confirm the experimental repeatability of the tests. 3.4 Commentary The observed failure modes for each pipe are shown in Table No 2. The mode and location of the failure were observed to be influenced by relative pipe sizes, soil density and end boundary condition of the branch pipe. Table 2 : The observed failure modes for different tests.

Test No. Observed failure modes 114D42-U 114D42-R 114L42-U 114L42-R

Crack formation observed at the tee. No damage was observed in branch and trunk

pipes [see Figure 5(a) and (b)].

114D16-U Branch pipe failed immediately after tee from buckling

114D16-R Branch pipe failed due to excessive yielding and leading to rupture

60D16-U 60D16-R

Very large strains observed in trunk line

As noted by the National Transportation Safety Board (NTSB, 1998), 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. In spite of this, NTSB (1998) reports, from field inspections, that most of PE pipe failures are attributable to brittle-like failures due to stress concentrations and not ductile failures, especially from a long-term performance point of view. In this context, assessment of the development of localized stress/strain concentrations is an important consideration, and the tests results from the present study provide an opportunity to examine this aspect for certain tee junctions. For example, as per Test No. 60D16-U, the presence of the tapping tee with a relatively larger branch pipe caused substantial anchoring of the trunk pipe at the tee (i.e., only small movements at the tee, see Figure 8), thus making the movement of the leading end of the trunk pipe essentially equal to the differential movement between the leading end and the tee junction. If strain localization was ignored, the axial strain of the pipe could be estimated simply using the measured string potentiometer displacements readings at the leading end and at the tee. The variation of axial strain along the pipe calculated on this basis is compared with those obtained from direct strain gauge measurements in Figure 9 (NOTE: comparison herein made at the instance where pipe movement at the leading end of the pipe is 100mm). It can be clearly seen that strain computations made using differential displacements alone are not only unsatisfactory but non-conservative.


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As indicated earlier, in some tests at a given cross-sectional location, strain gauges were placed at the invert of the pipe in addition to those at the crown. The measurements indicated that the strains at the pipe invert were smaller than those at the crown, again, confirming the need to account for the localized stress concentrations in the vicinity of the tee. Although it is not possible capture the maximum localized strain due to capacity limitations of strain gauges, values obtained from strain gauges within its measurable range seems still to provide a good lower bound approximation of the local strain magnitudes. For example, as noted in Figure 9, for a 100-mm pipe movement at the leading end of the pipe (displacement of 25 mm of the tee) would result in almost 6% localized strain in trunk pipe.

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Strain calculated from differential displacements

Figure 9. The strain distribution along the pipe from (a) strain gauge readings and (b) calculated from differential displacements of the string potentiometers, when leading end of the pipe moved by 100mm, [in 60D16-U tests] In all tests, a PVC protective sleeve was used at the joint to simulate the conditions of field installations [see figure 1(b)]. From the designers point of view, this sleeve is intended to provide additional bending stiffness at the tee and prevent shear (NTSB 1998). However, from the tests conducted herein, it was found that the sleeve can move along the branch pipe at larger displacements and may have very little reinforcing effect at the joint. It was also noted that if the sleeve is not properly locked in place, it will not contribute bending stiffness of the branch pipe. In 114*42-* tests, the effect of the sleeve was negligible since the weaker section of the “socket type” tapping tee was not protected by the sleeve [Figure 1 (b)]. It is also worth noting that there are no standard specifications for designing the diameter, length and stiffness of the PVC protective sleeve. (NTSB, 1998). A sleeve that is properly locked in place will likely contribute to the bending stiffness of the branch pipe.

4. CONCLUSION A full-scale testing program was undertaken to study the performance of several commonly used tee-junction configurations of buried PE piping (used in natural gas distribution systems) subjected to relative ground movements. The test results directly contribute to developing a reliable database for developing/validating numerical models to capture the response of pipeline distribution systems subjected to ground movements. The failure modes in the tested configurations were noted to be dependent on relative pipe sizes between trunk and branch pipes, density of soil backfill surrounding the pipe, and boundary conditions at the end of the branch pipe. The level of soil anchoring at the tee would govern the movement of the tee depending on the stiffness of the trunk pipe. The branch pipes were noted to be more vulnerable when the axial stiffness of the trunk pipe is more prominent. On the other hand, a branch pipe with a relatively larger diameter could cause significant soil anchoring at the tee; in turn, this could result in significant straining and failure of the trunk pipe. This confirms the importance of considering the possible restraining effects and branch pipe’s capacity to undergo bending and tension in numerical analysis. Special consideration also has to be given to the local strain capacity of the tee, particularly if the adjoining branch pipe is stiff. The results clearly highlight that strain computations made using differential displacements at certain pipe locations alone are not satisfactory and non-conservative. Although it is not possible to capture the very large localized strains, strain gauges would still provide a good lower bound approximation of the developed local strain magnitudes. Generally, the capacity of small pipes is difficult to determine due to its low stiffness, strength and high variability in field installation conditions. The results suggest that a properly designed PVC protective sleeve can improve the performance of the branch pipe. 5. ACKNOWLEDGEMENT The funding for this research provided by Terasen Gas Inc. is gratefully acknowledged. The authors also wish to extend their sincere thanks to Messrs. Doug Smith, John Wong, Bill Leung and Scott Jackson of Department of Civil Engineering for their technical assistance. The assistance given by undergraduate student Mr. Winson Cheung in performing the tests is also appreciated. References Anderson, C. (2004) Soil-Pipeline Interaction of Polyethylene

Natural Gas Pipelines in Sand, M.A.Sc Thesis, Dept of Civil Engineering, University of British Columbia.

Anderson, C., Wijewickreme, D., Ventura, C., and Mitchell, A. (2004) Full-scale Laboratory Testing of Buried Polyethylene Gas Distribution Pipelines Subject to Lateral Ground Displacements, Proceedings of 13th


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World Conference on Earthquake Engineering, Vancouver, B.C., August 1-6, Paper No. 1543

Gazagnaire, C. (2000) Long-term Performance of Buried High Density Polyethylene Plastic Piping, Master’s thesis, Florida Atlantic University, Boca Raton, Florida

NTSB. (1998) Brittle-like Cracking in Plastic Pipe for Gas Service, Special Investigation Report, National Transportation Safety Board, Washington, D.C.

PIPA (2001) Polyolefins Technical Information, Plastics Industry Pipe Association of Australia.

O’Rourke, T.D., Druschel, S.J., and Netravali, A.N. (1990) Shear Strength Characteristics of Sand-polymer Interfaces, Journal of Geotechnical Engineering, ASCE, 116 (3): 451-469

Stewart, H.E., Bilgin, O., O’Rourke, T.D. and Keeney, T.M. (1999) Technical Reference for Improved Design and Construction to Account for Thermal Loads in Plastic Gas Pipelines, Technical report, Cornell University, Ithaca, N.Y.

Wijewickreme, D., Sriskandakumar, S., Byrne, P.M. (2005) Cyclic Loading Response of Loose Air-pluviated Fraser River Sand for Validation of Numerical Models Simulating Centrifuge Tests, Canadian Geotechnical Journal, 42 (2): 550-561


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