Experimental Study of Stress Cracking in High Density Polyethylene Pipes. A Thesis. Submitted to the Faculty. Drexel University.

Transcription

1 Experimental Study of Stress Cracking in High Density Polyethylene Pipes A Thesis Submitted to the Faculty of Drexel University by Jingyu Zhang in partial fulfillment of the requirements for the degree of Doctor of Philosophy November 2005

2 ii Dedication To my family with love.

3 iii Acknowledgement It is a test of effort and persistence to complete a PhD research. I would like to take this opportunity to express my sincere appreciation to those who have assisted and supported me to make it a possibility during the last four years. First and foremost, I would like to thank my advisor Dr. Grace Hsuan for her guidance and advice. I also would like to extend my thanks to Dr. Robert Koerner; his advices and encouragement have helped me tremendously during the last four years. In addition, I would like to express my gratitude to Dr. George Koerner for his encouragement. I would like to thank deeply my colleagues and friends, Mr. Greg Hilley, Mr. Lei Lou, Mr. Songtao Liao, Mr. Mengjia Li, Mr. Sangsik Yeo, Ms. ShiQiong Tong, and Ms. Cynthia Baxindine for their help by performing the experiments, sharing experiences, and offering valuable discussions. The project is made possible by the support of the Florida State Department of Transportation. My appreciation also goes to the faculty in the Civil, Architectural, and Environmental Engineering Department at Drexel University for providing a great learning environment. Finally, I would like to thank my family in China for their everlasting love and support.

4 iv Table of Contents List of Tables viii List of Figures x Abstract.xiv 1. Background and Literature Review Introduction Pipe material HDPE pipe application Cracking of HDPE products Stress cracking mechanism Lustiger s microscopic model Crazing Environmental stress cracking Fatigue-related stress cracking Stress cracking test methods Residual stress Lifetime prediction methods Fracture mechanics method Shifting method Rate process method Stress Cracking in HDPE Protection Ducts in Segmental Bridges Introduction.32

5 v 2.2 Background Segmental bridges Tendon failures in segmental bridges Corrosion mechanism HDPE ducts included in this study Assessment of cracking mechanisms Introduction Macroscopic and microscopic evaluation Cracked samples from the MB Bridge Cracked samples from the SSK Bridge Summary of cracking mechanism evaluation HDPE duct properties Specifications for HDPE materials Specifications used for the ducts Test results of the HDPE ducts Correlation of SCR to other material properties of the ducts Recommended specification Assessment of stresses in the ducts Temperature-induced stresses Laboratory test model Results from test model FEM analysis Residual stresses 80

6 vi 2.7 Evaluation of fatigue failure Fatigue test Discussion of fatigue test results Summary Stress Cracking in Corrugated HDPE Pipes Introduction Background Corrugated HDPE pipes Failures in HDPE corrugated pipes Test materials SCR of liner Introduction Experimental design of liner test Data analysis Data analysis method Test results in water environment Test results in air environment Test results in Igepal environment Comparison of SCR in different environments Residual stresses Residual stress measurement Residual stresses effect on SCR Effect shown from four specimen configurations 109

7 vii Effect shown from comparison on liner and plaque specimens SCR of junction Introduction Experimental design of junction test Data analysis Comparison of deferent tests Comparison of liner and junction test Comparison of notched liner with liner and junction tests Lifetime prediction Data extrapolation methods Comparison of prediction methods Prediction using RPM Summary Conclusion and Future Work Summary HDPE ducts in segmental bridges Corrugated HDPE ducts Conclusion Future work..143 List of Reference..144 Appendix A: Data Analysis Method and Matlab Code Appendix B: Residual Stress Calculation 160 Vita...164

8 viii List of Tables 1.1 Comparison of test methods Information on samples retrieved from the bridges Specification designated in ASTM D Equivalency of ATM D3350 and D Original material specification for the HDPE ducts in each bridge Detailed original specification for the HDPE duct material Test results for the duct material properties Recommended specification for HDPE ducts Summary of the analytical model results Combined temperature-induced stress and residual stress Properties of the studied pipe samples Failure times of the four specimen configurations Test environments for liner test Summary of fitted curves of liner test on A Residual stress measurements and effect on failure times Summary of fitted curves of plaque and liner tests on A Activation energies from different tests Summary of fitted curves of junction test on A Comparison of the brittle region of liner and junction tests on A Comparison of acceleration effect of liner and junction tests..127

9 ix 3.11 The three constants used for RPM method Comparison of the two types of HDPE pipes...143

10 x List of Figures 1.1 Ductile failure in a HDPE pipe Brittle failure in a HDPE pipe Impact fracture surface under 1000x magnification SC fracture surface under 1000x magnification Failure modes that could occur in HDPE pipes Graphic illustration of ductile failure at molecular level Graphic illustration of brittle failure at molecular level Craze at the crack tip Test equipment for NCTL test Residual stress distribution caused by cooling Precast concrete box Construction of a segmental bridge Inside view of a segmental bridge Cross-section of a tendon in segmental bridges Corroded steel strands in segmental bridge tendons Close-up view of the corroded steel strands Corrosion mechanism of steel tendons Cracking in HDPE ducts Duct section that contains a full crack Duct section that contains part of the length of crack Drawing of sample MB 67-5-A.43

11 xi 2.12 Specimen 1 from sample MB 67-5-A Specimen 2 from sample MB 67-5-A Drawing of sample MB 38-4-A Specimen 1 from sample MB 38-4-A Specimen 2 from sample MB 38-4-A Specimen 3 from sample MB 38-4-A Specimen 4 from sample MB 38-4-A Drawing of sample MB A Specimen 1 from sample MB A Specimen 2 from sample MB A Specimen 3 from sample MB A Specimen 4 from sample MB A Drawing of column sample SSK 131-SB-SE Specimen 1 from sample SSK 131-SB-SE Specimen 2 from sample SSK 131-SB-SE Specimen 3 from sample SSK 131-SB-SE Correlation between SCR and MI Correlation between SCR and density Test model built by FAU Strain gage arrangements in the test model Void in the test model Strain gage measurement Two-dimensional FEM model of duct-grout system.77

12 xii 2.35 Mesh in analytical model First principal stress distribution when the void is 10% Void size effects on the principal stress Measurement of residual stress by slitting Combined temperature-induced stress and residual stress Fatigue test system Specimen in the fatigue test S-N curves on logarithmic scale for different duct samples Creep and stress relaxation behavior of HDPE pipes Cracking in a field corrugated HDPE pipe Typical fracture morphology from a cracked field corrugated HDPE pipe Geometry of type S corrugated pipe Liner specimen location Liner specimen A36 liner test at 60, 70, and 80 o C in water A36 liner test at 60, 70, and 80 o C in air A36 liner test at 50 o C in Igepal solution Compiled graph for A36 line test in all environments Activation energy calculation for A36 liner test in water at 600 psi Measurement of arc height of the specimen Comparison of test results from liner and junction test in Igepal solution Comparison of test results from liner and junction test in water Comparison of test results from liner and junction test in air..115

13 xiii 3.16 Junction configuration Junction test specimen location Typical configuration of junction specimen A24 junction test at 60, 70, and 80 o C in water The fitted curves and 97.5% lower confidence limits for junction test in water Comparison of test results at 60 o C in water on A Comparison of test results at 70 o C in water on A Comparison of test results at 80 o C in water on A Dependence of the ratio between junction failure time and liner failure time on stress and temperature Difference of activation energy in liner and junction tests Comparison of liner and junction tests at 80 o C on A Comparison of predictions for liner test in water on A Comparison of predictions for liner test in air on A Comparison of predictions for junction test in water on A Comparison of predictions for liner test in water on A A24 junction test data and prediction at 23 o C.134 B.1 Specimen configuration after annealing...160

14 xiv Abstract Experimental Study of Stress Cracking in High Density Polyethylene Pipes Jingyu Zhang Grace Hsuan, PhD Stress cracking (SC) is recognized as one of the major concerns for high density polyethylene (HDPE) pipes. SC is a brittle failure that occurs at a stress level lower than the short-term mechanical strength of the material. Many cases of SC have been reported in two types of HDPE pipes: HDPE ducts used in segmental bridges and corrugated HDPE drainage pipes. The causes of SC in these two types of pipes and their stress cracking resistance (SCR) properties are evaluated in this dissertation. Longitudinal cracking was observed in HDPE ducts and it was resulted by circumferential stress inducing from temperature cycles in the field and residual stress in the pipe. Conversely, the circumferential cracking in corrugated HDPE pipes was caused by longitudinal stress from bending and residual stress of the pipe. Majority of the fracture surfaces were covered by the fibril structure indicating that cracks propagated via a slow crack growth (SCG) mechanism. The study confirmed that the notched constant tensile load (NCTL) test can effectively distinguish the SCR of different HDPE ducts and the NCTL test is incorporated into the recommended material specification for quality control of HDPE ducts. Since fatigue lines were observed on the fracture surface, fatigue tests were adopted to estimate the lifetime of the duct under thermal cyclic loading in the field. For the corrugated HDPE pipes, the SCR evaluation focused on the finished pipe in order to incorporate processing effects. Two test methods were developed and

15 xv evaluated in the study, namely liner and junction tests. The liner test utilizes notched specimens to generate short and consistent failure times, and is good for QA/QC. On the other hand, the junction test challenges the junction where field cracking is observed. Thus, the junction data were used for predicting the long-term SCR of the pipe, and the reliable method was found to be the rate processing method (RPM).

16 1 Chapter 1: Background and Literature Review 1.1 Introduction The use of high density polyethylene (HDPE) pipes has increased significantly in last fifty years in different sectors of civil and environmental engineering. Generally, HDPE pipes have provided satisfactory performance; however, stress cracking (SC) was reported in various types of HDPE pipes. ASTM D883 defined SC as an external or internal rupture in a plastic caused by tensile stresses less than its short-term mechanical strength. According to the report by Hartt and Hsuan (2004), cracking was observed on some of the HDPE ducts in the segmental bridges located in the state of Florida. Another report by Hsuan and McGrath (1999) revealed many cases of cracking in corrugated HDPE drainage pipes. The causes of the SC in these two types of pipes are associated with material properties, geometry, manufacturing process, and field loading conditions. The objective of this study is to identify the cracking mechanism, develop effective test methods for evaluating stress crack resistance (SCR), and provide appropriate approaches to estimate the lifetime for both types of HDPE pipes. 1.2 Pipe material The conventional materials that are used for piping purposes are concrete and metal. The oldest material is vitrified clay; the newest materials are polymeric materials, in which HDPE is widely-used. Polyethylene has the simplest molecular structure of all polymers. It consists of two carbon and four hydrogen atoms in the basic repeating unit. Polyethylene is polymerized

17 2 from ethylene gas that is obtained from natural gas or crude oil. The polymerization conditions of low temperature, low pressure, appropriate catalysts (such as Ziegler-Natta catalyst), and co-monomers result in a linear polyethylene. Linearity indicates that there are limited branches in the polymer chains, so that the molecules can pack tightly. High density polyethylene (HDPE) is one type of linear polyethylene with a density range from to g/cc as per ASTM D883. HDPE exhibits high strength and modulus; thus, it is preferred for use in the manufacture of plastic pipes. Other common plastic materials used for piping include polyvinyl chloride (PVC), polypropylene (PP), polybutylene (PB) and acrylonitrile-butadiene-styrene (ABS). Along with HDPE, these thermoplastic pipes exhibit toughness, flexibility, high chemical resistance, light weight, easy installation, and low Manning coefficient, which make them suitable for engineering applications. 1.3 HDPE pipe application Plastic pipes have been used in pressure piping applications for many years. Sarkes and Smith (1983) pointed out that the use of plastic pipes began in the gas industry from In the early 1970s, plastic pipes started being used in highway drainage applications. Currently, HDPE drainage pipes are installed more frequently than all other plastic pipes combined. In the past thirty years, HDPE pipes have also been used as a protection layer for cables in segmental bridges, encasing steel strands and concrete to prevent corrosion. The three primary applications of HDPE pipes are described as follows.

18 3 Fuel pipes HDPE is now the dominant material used for fuel gas conveyance pipes. These pipes are smooth-walled, with relatively small standard dimension ratio (SDR), which is the ratio between the outside pipe diameter and the minimum pipe wall thickness. SDR is related to internal pressure capability, as shown in Equation (1.1): D t SDR 1 S = P = P (1.1) 2t 2 Where, S = circumferential stresses in the pipe wall, P = gas pressure, D = average outside diameter, t = minimum pipe wall thickness. The design involves keeping the working pressure in the pipe lower than the pressure rating of the pipe, which is defined by ASTM 2837 as the estimated maximum pressure that the medium in the pipe can exert continuously with a high degree of certainty that failure of the pipe will not occur. Due to the critical application of the gas pipe, the quality of the pipe is carefully evaluated. Significant amounts of research, funded by the Gas Research Institute, were carried out in the 1980s. The results of that research have greatly benefited the application of HDPE pipes in different fields. Drainage pipes Corrugated HDPE pipe is the most widely-used plastic pipe for gravity flow water systems, which include storm sewers, perforated under-drains, storm drains, slope drains, cross drains, culverts, and sanitary sewers. The diameters of these pipes span from 4 to 60 inches. In the majority of the applications, the pipes are buried underground. As such, they are designed to support the soil load and live load. The flexibility of the HDPE pipe

19 4 allows limited deformation, which will transfer a portion of the overburden onto the surrounding soil. Furthermore, because the stress relaxation is greater for the HDPE pipes than the soil, more overburden could be taken over by the soil with time. However, excessive deformation can lead to buckling of the pipes or/and jeopardizing of the stability of the pipe/soil structural system. Therefore, satisfactory performance requires HDPE pipes with sufficient stiffness. In order to increase the structural stiffness of the pipe and reduce the material cost, corrugation is incorporated into the pipe profile. According to AASHTO M294, there are three types of HDPE pipe profiles: Type C: This pipe shall have a full circular cross-section, with an annular corrugated surface both inside and outside. Type S: This pipe has a full circular dual-wall cross-section, with an outer corrugated pipe wall and a smooth inner liner. Type D: This pipe has a circular cross section consisting of an essentially smooth inner wall joined to an essentially smooth outer wall with annular or spiral connecting elements. Protection Ducts In this application, the HDPE pipes, called ducts, function as a moisture barrier. One particular example of this application is the sheathing of cables in segmental bridges. The HDPE duct surrounds the cement grout and steel to isolate them from the outside environment. The duct and grout then form a double corrosion-protection system for the steel strands. The ducts are smooth-walled pipes with diameters generally around 4 inches.

20 5 1.4 Cracking of HDPE products HDPE pipes have performed successfully in many applications; however, SC was found to be the cause of primary premature failure in all three types of applications described above. The Gas Research Institute published a report that summarizes many stress crack field failure cases of gas pipes and their forensic analysis (Gas Research Institute, 1984). Hsuan and McGrath (1998) investigated 19 stress crack field failure cases in corrugated HDPE pipes for highway applications. Recently, Hsuan and Hartt (2004) completed a report on the evaluation of SC of HDPE ducts in seven segmental bridges in the State of Florida. The cracking phenomenon raises concerns regarding the long-term integrity in some of the HDPE ducts. Recognizing the SC issue in the HDPE pipe, researchers have carried out extensive investigations. Their efforts and findings have contributed to the understanding of the cracking mechanisms of the HDPE material. According to Lustiger (1985), the cracking of HDPE pipes in the field can be categorized into three types: 1) third-party damage, 2) joint failure, and 3) material failure. The third-party damage is the result of improper construction practices. Joint failures are caused either by improper joining conditions, or by a material deficiency that inhibits proper fusion. Material failure is related to the polymer s inherent properties, poor pipe design and faulty manufacturing process. Ductile versus brittle failure The cracking of polyethylene pipes consists of two modes: ductile or brittle mode. According to Lu and Brown (1990), ductile failure is associated with macroscopic yielding. The time to failure of ductile failure is determined by creep rate. On the other

21 6 hand, brittle failure is associated with crack growth. Lu and Brown suggested that the two processes occur simultaneously; and the final failure depends on which process is faster under given stress, temperature and notch depth. Ductile failure Ductile failures exhibit large material pull-out (or yielding) adjacent to the failure location. The extreme example is the tensile test of the plastic samples. This type of failure requires relatively high applied stresses and failure takes place in a relatively short time. The mechanism is related to the viscoelastic behavior of HDPE materials and specifically refers to the creep rupture. The resulting failure shows large deformation accumulating in this process. Figure 1.1 shows the ductile failure in a gas pipe. Figure 1.1 Ductile failure of a HDPE Pipe Brittle failure The majority of cracking taking place in the field does not exhibit large deformation. There is no pull-out or thinning down of material adjacent to the crack. Figure 1.2 shows a field pipe with a brittle failure. This type of failure is defined as brittle (or brittle-like) failure.

22 7 Brittle failure in HDPE usually occurs under low stresses and takes a long period of time to propagate through the material s thickness via the process of slow crack growth (SCG). Lustiger (1987) stated that the SCG process can vary from hours to years at rates less than 0.1m/s. SCR is the least-desirable failure mode for HDPE products because it shows no sign prior to the failure. As a result, enormous attention is directed to the study of SCG. SCG can be resulted of creep- and fatigue-loading. Figure 1.2 Brittle failures in a HDPE pipe (From Hsuan, 1998) Brittle failure in HDPE pipes can also be caused by impact. The phenomena of SCG and impact facture tend to have similar failure appearance. However, they possess fundamentally different failure mechanisms. Lustiger (1985) compared the differences, which are listed as follows:

23 8 a. Impact failure happens fast, with the crack-growth rate close to the speed of sound (300m/s), called rapid crack propagation (RCP); SCG usually takes a relatively long time, ranging from minutes to decades (at a speed less than 0.1m/s). b. Impact failure tends to occur at lower temperatures and at high loads, whereas with increasing temperature at loads below yield strength, the tendency for SCG is increased. c. Impact fracture surfaces in PE display a flaky, scaly appearance (Figure 1.3), whereas SCG reveals a fibrous texture (Figure 1.4). Figure 1.5 summarizes the failures that could occur in HDPE pipes. In this study, the focus is on the phenomenon of SCG, which has been found to be responsible for many SC of HDPE pipes. Figure 1.3 Impact fracture surface under 1000x magnification (from Hsuan and McGrath, 1999)

24 9 Figure 1.4 SC fracture surface under 1000x magnification (from Hsuan and McGrath, 1999) Figure 1.5 Failure modes that could occur in HDPE pipes

25 Stress cracking mechanism Understanding the failure mechanism is a key component to improving the SCR of HDPE pipes. This section describes the stress cracking process based on macroscopic and microscopic perspectives Lustiger s microscopic model The microscopic aspect of the stress cracking mechanism is not yet fully understood. In 1985, Lustiger proposed a simple model to explain the microscopic deformation in ductile and brittle failures. The HDPE material comprises an ordered crystalline region and a random amorphous region. The crystalline region consists of packs of folded molecules named lamella, which are separated by the amorphous region. The intercrystalline polymer chains play an important role in the deformation. There are three types of intercrystalline chains: Cilia chains suspended from the end of a crystalline chain Loose loops chains that begin and end in the same lamella Tie molecules chains that begin and end in adjacent lamellae When tensile load is applied normally to the face of the lamellae, the tie molecules are pulled and deformed. Since the tie molecules are intricately tangled, they can be viewed as reinforcing elements. This model is sometimes called the mortar and brick model, in which lamellae may be viewed as bricks and the tie molecules as mortar. The mortar holds the bricks together. At high stresses, the tie molecules are pulled until they cannot

26 11 support the applied stresses. Then the lamellae break up into smaller units, as illustrated in Figure 1.6. In this case, ductile failure happens. Figure 1.6 Graphic illustration of ductile failure at the molecular level (From Lustiger, 1985) When the stress level is low, tie molecules can gradually disentangle and relax with time. As a result, interlamellar failure will occur. The process is shown in Figure 1.7. This failure yields a brittle-like fracture surface Figure 1.7 Graphic illustration of brittle failure at the molecular level (From Lustiger, 1985)

27 12 The key component of this model is the role of tie molecules during the fracture process. Because tie molecules bridge adjacent lamellae, their density, integrity and ability to remain entangled are critical. Assuming the abilities of all the tie molecules are the same, the model suggests that materials with fewer tie molecules are more susceptible to brittle failure than those with greater number of tie molecules. However, if the density of tie molecules is too high, it is usually at the expense of the material stiffness. Some of the parameters that influence the number of tie molecules in the HDPE follow: Molecular weight: Higher molecular weight indicates longer polymer chains, which can result in more tie molecules and more effective tie molecule entanglements. Comonomer content: Polyethylene is a product of the copolymerization of ethylene and comonomers. The comonomers form short branches along the linear polyethylene molecules that tend to inhibit crystallization. A higher comonomer concentration leads to more tie molecules. Furthermore, the branches contained in the chains inhibit the ability of the tie molecules to slip past one another Crazing Crazing can be considered as the macroscopic aspect of stress cracking. The craze, or damage zone, is the region ahead of the crack tip which consists of voids and stretched fibrils. The structure of the craze is illustrated in Figure 1.8. Brown et al. (1985) suggested that the brittle failure process can be categorized into a series of events. First, a craze is formed at the root of the notch immediately following the application of

28 13 loading. At the root of the craze, a plastic zone is generated due to the localized yielding of the material. Crack tip Micro-fibrils Yielding Figure 1.8 Craze at the crack tip The craze remains stable, with the micro-fibrils sustaining the stresses. As time passes, the craze grows slowly by stretching the micro-fibrils. The rupture of the micro-fibrils near the base of the craze leads to a growing crack. When the remaining ligament reaches the critical size, complete failure occurs. Chudnovisky et al. (2003) suggested that crack propagation can be viewed as a step process. They named the craze itself the process zone (PZ), and the craze and crack together the crack layer (CL). Under tensile stresses, the fibers inside the PZ creep. After a period of time, the fibers break down, and the crack extends into the PZ. The time until the PZ breaks down corresponds to the arrest time in the crack propagation process. After that, a new PZ is developed from the

29 14 intact material ahead of the crack. The newly-created PZ stops the cracking process while the micro-fibrils undergo creep deformation. The process repeats itself. The time to formation of a new PZ is significantly shorter than the arrest time. This process of steps will end when the crack becomes unstable. Most literature divides the SC process into crack initiation and crack growth. The initiation period is the stage before the craze is formed. Once the craze is formed, the crack propagation begins. Cracks initiate from defects in the material such as flaws, notches created by installation, impingement, and bending loads. In the laboratory tests, notched specimens were used to generate consistent failure time. The failure time represents time for both crack initiation and propagation (Cassady and Uralil, 1985). Bragaw (1980) indicated that for HDPE pipes, one-third of the time to failure is consumed by crack initiation. Others suggested that the initiation of the cracking can represent as much as 90% of the total cracking process. Bragaw (1983) stated that the initiation time becomes infinite when the stress is below a certain threshold Environmental stress cracking Environmental stress cracking (ESC) occurs when PE is subjected to stresses in the presence of various environmental agents. The fundamental molecular mechanism of ESC in PE is still debated. Nevertheless, ESC and SCG share many similarities, such as load and temperature dependence of failure time, and brittle-like failure surface. Therefore, it is thought that they probably have a common molecular deformation mechanism. If so, it would be valuable to use ESC as a tool to evaluate the long-term

30 15 behavior of polyethylene material, since the ESC process takes a much shorter time to complete than conventional SC tests. The acceleration effects of environmental agents in SC have been studied by many researchers. Hopkins et al. (1950) suggested that the surfactant exerts a spreading pressure on intrinsic flaws and cracks on the polymer surface. The pressure works with the applied stress and results in crack initiation. Isaksen et al. (1963) suggested that ESC occurs because the agent is absorbed preferentially by the most highly stressed crystallites, which act as a stress concentration raiser. Some researchers believe that the environmental agent induces the plasticization of tie molecules and enhances their disentanglement, subsequently accelerating the SC process Fatigue-related stress cracking While previous sections discuss cracking under static loading conditions (creep), dynamic loading (fatigue) can also induce cracking. It has been observed that dynamic loading can significantly increase crack propagation. Nishimura and Shishich (1985) found that fatigue testing can shorten the creep failure time by more than two orders of magnitude, while Shah et al. (1998) indicated acceleration of up to three times in the fatigue test. Zhou and Brown (1993) found that the fastest failure in the fatigue test occurs when the loading is in tension-compression mode, and is probably due to the buckling of the fibrils under the compressive load. Parsons et al. (1999) found that the size of craze is controlled by mean stress alone, and the crack growth rate is related to both the maximum and mean stress through a power-law relationship.

31 16 Since fatigue loading can significantly accelerate the cracking rate, the test can be conducted at room temperature while still achieving reasonably short testing time. However, questions arise regarding the similarities between fatigue and SCR. Some of the similarities include: 1) their fracture surfaces have a similar appearance; 2) the fatigue crack growth exhibits step propagation, i.e., a craze is formed at the crack tip, which relieves the stress, and the crack growth is stopped as long as the craze stays stabilized; 3) correlation between creep and fatigue has been observed. Zhou et al. (1989) performed a series of fatigue and creep tests and found that there is a linear relationship on a log-log scale between the cycles to failure under a fatigue test and the time to failure under a constant load test. Their finding suggested that fatigue test can be used to predict the creep fracture of polyethylene. Nevertheless, current practice of fatigue test is limited to material rating purpose only. 1.6 Stress cracking test methods The cracking of HDPE pipes can take place as early as a few months after installation, but the majority of the cracking failure occurs after years. This makes it impractical to study the SC by simulating the service condition in the laboratory. Thus, a variety of acceleration methods have been developed to target a reasonably short testing time (or failure time). The common approaches to achieving this objective include introducing a stress concentration (e.g. notching), and utilizing elevated temperatures, environmental solutions, high stress, or fatigue loading. A brief introduction of various test methods used to evaluate the SC of HDPE materials is presented in this section.

32 17 Bent strip test (ASTM D1693) This test was the dominant QA/QC test for polyethylene materials in the 60s and 70s. It involves cutting ten rectangular specimens notched on the surface longitudinally. The specimens are then bent into a 180 o arc and confined within the flanges of a small metal channel. The entire assembly, with notched specimens, is immersed in either 10% or 100% Igepal solution at temperatures of 50 o C or 100 o C. There are three test conditions, defined by the test specimen size, notch depth, and test temperature. The specimens are examined after certain periods of time and the percentage of the failed specimens is recorded. The test duration varies from 24 hours to 1000 hours, depending on the specification set by different HDPE industries. The bent strip test is simple and easy to perform. However, significant stress relaxation occurs during the test, and rate of stress relaxation is difficult to quantify. This test has been found to be insufficient to distinguish the SCR property of current HDPE materials; furthermore, the test results contains a large standard deviation. Pennsylvania notched test (ASTM D1473) The Pennsylvania notched test (PENT) was developed by Dr. Norman Brown at the University of Pennsylvania, (Brown et al., 1989). Test specimens can be taken from compression molded plaques or manufactured pipes. The compression molded plaques are made according to ASTM D4703, with some modifications. After the resin is heated to the set temperature, pressure is applied and removed several times to eliminate voids. The plaque is then slowly cooled to room temperature to achieve a high crystallinity and to minimize residual stresses. Specimens of 10x25x50mm (0.4x1x2 inches) bars are taken from the plaque. Two side notches of 1mm deep and one side main notch of

33 18 3.5mm (0.14 inches) are produced in the same plane. The two side notches are referred to as side grooves, which are used to promote the plane strain condition. The main notch is where the cracking is expected to initiate. The specimens from manufactured pipes are prepared based on the pipe diameters. For pipes larger than 25mm (1inch) in diameter, rectangular bars are cut from the pipe walls; for those with diameters less than 25mm (1 inch), sections of the pipe are used for the test. The notches on the specimens are made according to the pipe wall thickness. The specimens are tested at 80 o C in air under a single stress of 2.4 MPa (348 psi). Full notch creep test (ISO16770) The full notch creep test (FNCT) was developed by Nishio et al. (1982), and is the preferred test method in Europe due to its shorter failure time compared to the PENT test. The FNCT specimen is a square section of 10x10mm (0.4x0.4 inch) bar (about 3.5 inches long) with four coplanar notches 1.5mm (0.06 inch) deep, made by a razor. The test is performed in a liquid environment at 80 o C under a single stress level. One common choice for the liquid is the 2% Igepal CO-630 mixed in 98% demineralized water. The test is suitable to evaluate the latest pipe resins, such as PE100, because it takes a few hundred hours to fail the specimens, compared to a few thousand hours for the PENT test. Notched constant ligament stress test (ASTM F2136) The notched constant ligament stress (NCLS) test was developed from the notched constant tensile load (NCTL) test and is designed to evaluate corrugated HDPE pipes and pipe resins. The test uses dumbbell-shaped specimens with notch on one side of the specimen s surface. The depth of the notch is 20% thickness of the specimen. The

34 19 major difference between this and other tests is that the specimens are taken from compression molded plaques in the NCLS test, instead of the end products in the NCTL test. Also, a single applied stress of 600 psi is used in the NCLS test. The test is performed in 10% Igepal solution at 50 o C. Figure 1.9 shows the test equipment. Figure 1.9 Test equipment for NCLS test Hydrostatic stress rupture test (ASTM D1598) This is a performance test and is used for pressure pipes. A section of manufactured pipe with caps on both ends is subjected to a constant internal water pressure at a desired test temperature. The test continues until the pipe fails. Even with elevated test temperatures and high stresses, this test takes a long time to finish. The crack initiation time based on material defects themselves occupied a significant part of the testing. The test truly reflects the performance of the pipe under the designed pressure; however, a large scatter in the test data.

35 20 Notched pipe test (ISO 13479) The potential scatter in the hydrostatic rupture can be avoided by introducing notches in the pipes. In the United Kingdom, both gas and water industries have adopted the notched pipe test (NPT) for quality assessment. This is an elevated temperature stress rupture test on a pipe with machined axial V notches. Four notches are equally-spaced around the circumference of the pipe. The notches run along the longitudinal axis of the pipe, with length equal to the pipe s outside diameter, and the depth is to give a remaining ligament of 80%±2% of the minimum wall thickness of the pipe. The pipe is tested at a stress level of 4 MPa (580 psi) for PE80 and at 4.6 MPa (667 psi) for PE100 materials. It is recommended that at least three tests be performed with no failure before 165 hours for both PE80 and PE100. The test environment is air at 80 o C. Fatigue test All of the above-described test methods are creep-related. As discussed earlier, fatigue testing is an alternative method that accelerates crack growth by applying cyclic stress. The test is best performed at a high frequency in order to decrease test time. However, increasing frequency can induce heat effect in the specimen. Most fatigue tests for polymers use frequencies around 1 Herz. The specimen geometry and stresses vary from study to study, depending on the test material. However, notched specimens have been commonly used to accelerate the crack initiation time. The fatigue cycles could be controlled by either load or strain. Table 1.1 summarizes the SC test methods.

36 21

37 Residual stress Aside from externally- applied stresses, residual stress can also be a stress factor causing SC. It is known that all plastic pipes contain residual stress introduced during the manufacture processes. The magnitude and distribution of residual stress vary significantly depending on manufacturing process. One of the main factors causing residual stresses is differential shrinkage through the pipe wall during the cooling process. Figure 1.10 shows the residual stress generated from two cooling approaches. Figure 1.10 Residual stress distribution caused by cooling Maxwell (2001) suggested that if cooling takes place from only one side of the wall, the cooled side will shrink rapidly. The newly-solidified section constrains its adjacent part from shrinking freely. As a result, compressive residual stress is created in the cool side and tensile residual stress is formed in the opposite side (Fig. 1.10(a)). However, if both sides are cooled simultaneously, compressive stress will be created near the two surfaces and tensile stress in the center (Fig. 1.10(b)).

38 23 Residual stresses can also be introduced by the drawing of the polymer chains during the manufacturing process. Wong (1983) found that biaxial tensile stresses are presented on the inner wall and biaxial compressive stresses on the outer wall for certain types of smooth pipe. Williams (1993) measured the residual stresses in both longitudinal and circumferential directions of extruded isotropic smooth pipes, and found the tensile residual stresses ranging from 2.12 to 4.07 MPa (307 to 590 psi). For anisotropic drawn pipes, residual stresses on the inner surface varied significantly, from tensile stress of 0.25 MPa (36 psi) to compressive stress of 2.91 MPa (422 psi). Chaoui and Moet (1987) studied the distribution of residual stress through the smooth pipe wall thickness and found that 24% of the inner wall consisted of tensile residual stresses, while the remaining thickness had compressive residual stresses, with the maximum value at the outer surface of the pipe wall. Kanninen et al. (1993) found that the distribution of circumferential residual stresses exhibited a parabolic shape through the smooth pipe wall, with tensile stress in the inner wall. High residual stress in the pipe could have significant impact on the SC property of the pipe. Therefore, quantifying the residual stresses in the pipe is essential in predicting the long-term behavior of the pipes. In this study, residual stresses in the smooth HDPE duct and corrugated HDPE pipe are evaluated. Tremendous efforts have been made to understand the residual stresses in metal pipes. Conversely, the evaluation of residual stress in plastic pipes is very limited. Many

39 24 methods used to assess the residual stress in plastic pipes were adopted from those used for metal pipes. Some of these methods are presented here. Slitting and parting-out method In order to measure residual stress in the circumferential direction of the pipe, a ring specimen with width of one-inch is removed from the pipe. The ring is then slit open. If the pipe has tensile stresses on the inside wall, the ring will tend to close and decrease the diameter. By measuring the changing diameter over time, the maximum residual stress can be calculated. For measuring the longitudinal residual stress, a small rectangular shaped section is cut into the end of the pipe. The displacement at the end of the removed section is monitored with time. Layer removal method The layer removal method is the most widely-used technique to measure residual stresses. A rectangular specimen is taken from the region of interest in the pipe, and then layers are removed from one side of the surfaces. The specimen undergoes dimensional change after removal of each layer. The average stress in each layer is calculated based on the dimensional change. A plot of the residual stress across the thickness can be obtained. However, the machine sectioning can introduce undesirable stresses to the test sample that can be either surface residual stresses or gross yielding stresses, reducing the accuracy of the measurement.

40 25 Hole drilling method (ASTM E837) This method involves drilling a hole into the material and measuring the surface strains in the vicinity of the hole during the drilling. Electric resistance gages are usually used to determine quantitatively the strains close to the drilling. The strains are used to calculate the biaxial stresses and their distribution at and near the surface of the part. There are a number of limitations to this method: the test sample must be wide compared with the diameter of the drilled hole; the residual stresses must be constant in the drilling area; inelastic flow should not occur during and after the drilling; and the method can measure only the residual stresses on the surface or at very limited depth. Annealing method This method is based on the assumption that a temperature increase does not change the molecular structure of the material significantly. The operation involves taking the specimens from the test section of the pipe and heating them to a suitable temperature. After a certain amount of annealing, the residual stress is thought to be completely relieved. The residual stresses are then measured according to the change of dimensions of the specimen prior to and after the annealing. There are other test methods, such as X-ray diffraction, ultrasonic method, indentation, and stress corrosion method. They are either still under development or used less often. 1.8 Lifetime prediction methods All of the SC tests described here use high temperatures to accelerate the failure process. Prediction methods are required to extrapolate the test data to the service temperature.

41 26 These methods can be categorized into three groups: the fracture mechanics method, the shifting method, and the rate process method (RPM) Fracture mechanics method Fracture mechanic theories can be used to model the failure time of SC. Liner elastic fracture mechanics (LEFM) have often been used to analyze the cracking of pipes. Numerous studies have been conducted in an effort to explore the relationship between the stress intensity factor (K) and the crack growth rate. Some semi-empirical models have been developed; however, the reliability of these models is not certain. Chan and Williams (1983) found that the steady state of the crack growth rate assumed by most of the fracture mechanics models exists only when the growth rate is at and above 10-9 m/sec. However, the actual crack growth rate in field application is well below that value. Popelar and Staab (1983) state that LEFM frequently fails to predict the cracking behavior of polyethylene pipes because it cannot properly account for the creep effect. Due to the semicrystalline nature of HDPE, the validity of linear elastic fracture mechanics (LEFM) is still controversial. More appropriate fracture mechanics techniques are needed to analyze cracking in plastic piping. The J-integral method has been extensively studied; however, the complexity of the method keeps the model from practical application Shifting method Boltzmann (1872) developed the fundamental equation for linear viscoelasticity, and one of the applications is the time temperature superposition (TTS) principle. By measuring a mechanical property at a series of temperatures, a master curve at a targeted temperature

42 27 can be obtained through shifting of the curve at each test temperature. The master curve can span decades of time, from which the material s lifetime at the targeted temperature can be predicted. Any individual test temperature, T, and the targeted temperature T r, have the following relationship: D ( t, T ) = D ( t / a T, T r ) (1.2) Where a T is the time-temperature shift factor, and is given by the Williams-Landel-Ferry (WLF) equation: loga T C1 ( T Tr ) = (1.3) C + ( T + T ) 2 r Where, C 1 and C 2 are constants. The assumption of TTS is that the material structure does not change during the test. Faucher (1959) showed that TTS is only valid in the linear deformation range below the material s melting temperature. He believed that when the temperature is above the melting point, the crystallinity in the PE will change. Onogi et al. (1967) suggested that the temperature affects the mobility of the crystallites. Other researchers (Popelar et al. 1990; Thomas, 1997) suggested that the high temperature leads to recrystallization, in which small crystallites grow into large crystallites. Popelar et al. (1990) found that horizontal shifting (temperature) alone cannot achieve a coherent master curve, due to the effect of temperature on crystallinity. However, with the help of vertical shifting, a master curve can be obtained. Since the method involves both the horizontal and vertical shifting of the curves, it is called bi-directional shifting. Popelar et al. (1990) analyzed a

43 28 large number of laboratory test data on relaxation moduli, stress rupture, and slow crack growth, and found that horizontal and vertical shift functions are universal for MDPE and HDPE used for gas pipes. The expressions for the shift factors are a b T T = exp[ 0.109( T T )] (1.4) r = exp[ ( T T )] (1.5) r Where a = horizontal shift factor; b = vertical shift factor; T T T = lab test temperature; T r = arbitrary reference temperature (or targeted temperature). Lu and Brown (1991) also proposed another shifting equation. These equations are based on the Arrhenius equation discussed in the next section. Q 1 1 h = 0.43 R (1.6) T T r 1 v = C T 1 T r (1.7) Where h = horizontal shift factor; v = vertical shift factor; Q = activation energy; R = gas constant; C = coefficient Rate process method (RPM) The Swedish chemist Arrhenius found empirically that the logarithm of the chemical reaction rate varies as the reciprocal of absolute temperature, provided that the range of the temperature does not effect large structure change in the material. It can be expressed in Equations 1.8 and 1.9. k E / RT = k0e (1.8)

44 29 Or ln k = ln k E / RT (1.9) 0 Where k = kinetic rate constant; k 0 = pre-exponential kinetic rate constant; E = apparent activation energy; R = universal gas constant, which is 8.314J/mole; T = absolute temperature. Equation 1.9 can be simplified and used to describe the relation between failure time, t and temperature, T: B log t = A + (1.10) T The interested failure time at a certain temperature can be obtained by the transformation of the Arrhenius equation: log t t E 1 1 = ( ) (1.11) R T r T r Where t= failure time at test temperatue ; t r = failure time at reference temperature; T= test temperature; T r = reference temperature; A, B = constants The Arrhenius equation provides the theoretical basis bridging the time to failure at different temperatures. The fundamental theory for the rate process method (RPM) is the Arrhenius equation. The development of RPM is illustrated as follows: It has been found that failure time and applied stress have a linear relationship on log-log axis logt = A + B logσ (1.12)

45 30 Where σ = applied stress; t= corresponding failure time Combinig Equations 1.10 and 1.12, we arrive at the RPM model introduced by Bragaw (1983): B logt = A + + C logσ (1.13) T A similar approach has been used to develop other RPM mathematical models, as follows. Three coefficient models include: B C logσ logt = A + + (1.14) T T B C σ log t = A + + (1.15) T T Four coefficient models : B C logσ log t = A D T (1.16) T T B C logσ logt = A D T logσ (1.17) T T B C logσ logt = A D logσ (1.18) T T Six coefficients model : logt = 1 B + E ( C + F) logσ 1 B E ( C F) logσ [ A + D + + ] [ A D + + ] 2 T T 2 T T (1.19) Where A, B, C, D, E, F = constants

46 31 Eugene et al. (1985) discussed the pros and cons of these models. It is found that the four coefficient models lead to better lack-of-fit values, which means the highest probability for regression line extrapolation. However, the three coefficient model can provide good fitting as well, without the complexity of a fourth term. Furthermore, while adding more terms may improve the fit, it also increases the uncertainty of the predictions. Therefore, Equation (1.14) is the model recognized as the most suitable in predicting the failure time, and is adopted by both ASTM and ISO.

47 32 Chapter 2: Stress Cracking of HDPE Ducts in Segmental Bridges 2.1. Introduction This chapter focuses on the SCR of the HDPE pipes that serve as a protective layer for post-tensioning tendons in segmental bridges. These pipes are commonly called HDPE ducts. Post-tensioned tendons are a key structural component of the segmental bridge, holding the precast concrete boxes together. The tendons consist of three parts: multiple high-strength steel strands, cement grout fill, and HDPE ducts. The cement grout and the HDPE duct serve as double protection layers for the steel strands encased inside. However, cracking of HDPE ducts can lead to corrosion of the steel strands by allowing moisture to penetrate the double protection layers. Cracking of the ducts should, therefore, be prevented. A recent survey on post-tensioned tendons (Hartt and Hsuan, 2004) reported that cracking appeared in the HDPE ducts of several segmental bridges in the State of Florida. A project was therefore initiated by the state of Florida to investigate the causes of the duct cracking. The project was carried out by Drexel University and Florida Atlantic University (FAU). This chapter consists of two parts, each covering one part of the investigative project. During part one of the project, Drexel University investigated the field-cracked HDPE ducts by identifying the cracking mechanisms, selecting appropriate test methods to assess the stress-crack resistance of HDPE ducts and recommending specifications to improve the quality of HDPE ducts. Part two (performed by both FAU and Drexel University) concentrated on the evaluation of stresses in the HDPE duct, together with service life assessment by fatigue test.

48 Background Segmental bridges Construction of precast segmental bridges is a relatively new technology. The technology was invented in France in the early 60s, and used widely in Europe during the late 60s and 70s. The first segmental bridge constructed in the United States (US) was the John F. Kennedy Memorial Causeway near Corpus Christi, Texas, in In the 80s, segmental bridges were constructed in many states. A precast segmental bridge is an assembly of precast concrete members, which are manufactured in a concrete plant, often near the construction site. First, the piers are built by stacking one precast block on top of the other. The segments of concrete boxes (Figure 2.1) are then hoisted and lowered on top of the piers to form the bridge deck. Figure 2.2 shows a segmental bridge under construction. Figure 2.3 is a picture taken inside the box, showing the tendons. The steel tendons are designed to connect the box girders and reinforce the structure. There are two types of tendons: internal and external. The internal tendons are cast inside the concrete box and cannot be inspected. The external tendons are outside the concrete, as shown in Figure 2.3. The construction method for external tendons involves housing the steel strands in the HDPE duct. The steel strands are gripped at both ends, pulled and anchored. Cement grout is then introduced into the duct, filling up the space and encasing the tensioned steel strands. Compared with conventional bridges, precast segmental bridges are more economical (especially when covering large spans), require less construction time and easy

49 34 maintenance, provide improved durability and appealing aesthetics. In 1999, the American Segmental Bridge Institute (ASBI) performed a comprehensive survey, and found satisfactory performance in 99% concrete segmental bridges. Figure 2.1 Precast concrete box Figure 2.2 Construction of a segmental bridge

50 35 Tendons Figure 2.3 Inside view of a segmental bridge Tendon failures in segmental bridges The integrity of the tendons is key for the safety of segmental bridges, since they are the reinforcing elements that hold the superstructure together. Therefore, corrosion of the steel strands must be prevented. The HDPE duct and cement grout act as double protection layers for the encased steel strands. Figure 2.4 shows the cross-section of a typical tendon. Figure 2.4 Nomenclature for cross-section of a segmental bridge tendon

51 36 Two corrosion-induced failures occurred in England: the Bickton Meadows footbridge in 1967 and the Ynys-Y-Gwas Bridge in The failures resulted in a ban on the use of post-tensioned bridges in the United Kingdom from 1992 to In the US, during the spring of 1999, a corrosion-related failure of an external tendon was found in the Niles Channel Bridge, Florida, after the bridge had seen 16 years of service (Powers, 1999). Niles Channel Bridge is one of a series of low-level segmental bridges stretching over seawater in the Keys area. Further inspection revealed that two steel strands were corroded in the tendon anchorage. In 2000, due to corrosion problems, eleven tendons out of a total of 846 were replaced in the Mid Bay Bridge after seven years of service. Also, in the same year, numerous corroded steel strands were discovered in segmental piers of the Sunshine Skyway Bridge, built in The corrosion was a result of seawater entering the ducts through the split in the ducts. Figure 2.5 and 2.6 show the corroded steel strands. Figure 2.5 Corroded steel strands in segmental bridge tendons (Mid Bay Bridge)

52 37 Figure 2.6 Close-up view of the corroded steel strands General Corrosion mechanism In good-quality grout, the encased steel strands are protected against corrosion by the alkalinity of the cement paste, whose ph is typically between 12 and 14. In a highlyalkaline environment, the surface of the steel is passivated and protected by the formation of an oxide film. The passive oxide film is stable at ph values greater than approximately 9.5 in a chloride-free environment. The ph value that is necessary to maintain the passive oxide film increases with increasing chloride content. The initiation of corrosion of the steel requires the breaking down of this oxide film, and this can occur by either diffusion of chlorides into the concrete or carbonation of the concrete. Carbonation is the result of a reaction of CO 2 in the air with Ca(OH) 2 in the concrete. As a result, the ph of the concrete gradually decreases. When the ph value of the concrete falls below 9.5 (in a chloride-free environment), the oxide film starts breaking down. The corrosion process is illustrated in Figure 2.7..

53 38 Figure 2.7 Corrosion mechanisms of steel tendons (From Hamilton et al., 1995) Tendons are in greater danger in bridges that are built above or near seawater, due to the high chloride content of sea water. Furthermore, the potential for steel corrosion is greater when the protective HDPE duct cracks. Thus, the integrity of the HDPE ducts must be maintained throughout the service life of the bridge HDPE ducts included in this study In this study, the HDPE ducts of seven bridges in the state of Florida were evaluated. Samples were retrieved from both cracked and uncracked ducts. Information on the bridges and the number of retrieved field samples is shown in Table 2.1. The ducts in MB and SSK bridges had significant cracking problems.

54 39 Table 2.1 Information on samples retrieved from the bridges Bridge Total No. of No. of County Cracking Field Sample Field Cracked Location Condition Age (yr.) Samples Samples Mid Bay Severe Okaloosa (MB) Cracking Garcon Point Santa Rosa No Cracking Seven Mile Monroe No Cracking Skyway (SSK) Channel Five Long Key Niles Channel Pinellas Monroe Monroe Monroe Severe Cracking No Cracking No Cracking No Cracking column 4 34 span The seven bridges are located in different regions of the state. The MB Bridge is situated in the northern part of Florida, where maximal ambient temperatures range from -10 to 40 o C. Seven Mile, Channel Five, Long Key, and Niles Channel bridges are located in the Keys region, with maximal temperatures ranging from 2 to 38 o C. Garcon Point, Skyway bridges are located in the central part of the state, with temperatures ranging from -5 to 40 o C Assessment of cracking mechanism Introduction The cracked ducts are coded with respect to their positions in the bridge. Figure 2.8 shows a cracked duct. The cracked section of the duct was removed from the tendon according to a specific sampling procedure so that the fracture surface of the crack would

55 40 be protected. Figures 2.9 and 2.10 show the sampling protocol for obtaining a full crack or part of the crack, respectively. Figure 2.8 Cracking in HDPE ducts 1 st Cut 2 nd Cut 2 nd Cut ~ 4 feet 1 st Cut Crack Figure 2.9 Duct section that contains a full crack 2 nd Cut 1 st Cut 2 nd Cut ~ 4 feet Crack Figure 2.10 Duct section that contains part of the length of the crack

56 41 The cracking mechanisms were assessed by both macroscopic and microscopic examinations of the cracks. Once the mechanism is identified, the appropriate test methods can be employed to assess the SC properties. Furthermore, the cause of the crack initiation can be identified by the microscopic examination Macroscopic and microscopic evaluation The retrieved field samples were shipped to the laboratory for examination. Each cracked duct was photographed and sketches of the cracks were drawn. Specimens for microscopic examination were then taken at various locations along the cracks. The fracture surfaces of the specimens were examined under a scanning electron microscope (SEM). Three cracked samples from MB and one cracked sample from SSK were selected for presentation Cracked samples from the MB Bridge The MB Bridge presented the most severe cracking problem in the HDPE ducts. Three samples retrieved from the MB Bridge were chosen for examination: 67-5-A, 38-4-A, and A. Sample MB 67-5-A MB 67-5-A (Figure 2.11) features a short crack oriented longitudinally along the duct. The crack is located near the 8 o clock position of the duct circumference. The crack length on the inner surface of the pipe is about 5 inches, and the crack length on the outer surface of the pipe about 4.5 inches. Specimens 1 and 2 are located near the crack tips. Figure 2.12(a) shows the general view of the fracture surface of Specimen 1. The detailed fracture surface under 1000x magnification illustrates the fibril structure (Figure

57 (b)). This is an indicator of SCG mechanism. The small size of these fibers suggests that the applied stress is relatively low during crack propagation. There are many impurities observed on the fracture surface, as seen in Figure 2.12 (c). These impurities may be the origin of the crack initiation.

58 43 Figure 2.11 Drawing of sample MB 67-5-A Figure 2.12 Specimen 1 from sample MB 67-5-A: (a) general view of fracture surface; (b) fibril fracture morphology in area A ; (c) a detailed view of an impurity.

59 44 Figure 2.13 shows the fracture morphologies of the specimen, which are very similar to those in Figure Small fibrils on the fracture surface in Figure 2.13 (b) indicate the slow crack growth mechanism. Again, impurities (Figure 2.13 (c)) are found in areas B and C, and these could initiate the crack. Figure 2.13 Specimen 2 from sample MB 67-5-A: (a) general view of fracture surface; (b) fibril fracture morphology; (c) a detailed view of an impurity in area B and C.

60 45 Sample MB 38-4-A In sample MB 38-4-A (Figure 2.14), two cracks can be observed, Crack B1 and B2. Both of them occurred at the 12 o clock position. They run along an irregular line and overlap each other by about 1 inch at the end without connecting. Crack B1 is about 13 long on both the inner and outer surfaces, and B2 is about 34 inches on both surfaces. Figure 2.14 Drawing of sample MB 38-4-A Four specimens were taken at different positions on the crack in sample MB 38-4-A. Specimen 1 is located near the end of Crack B1, where the Crack B1 and B2 join. Specimens 2, 3, and 4 were taken from Crack B2 and located in the right, middle and left section of the crack, respectively. The fracture morphology of Specimen 1 is shown in Figure The overall microstructure on the fracture surface is fibril structure, as can be seen in Figure 2.15(b). In this duct, air bubbles can be observed. No clear crack initiation point can be determined. In the upper half of the fracture, a horizontal line can be identified. This is believed to be a sign of the crack arresting during a fatigue process.

61 46 Figure 2.15 Specimen 1 from sample MB 38-4-A: (a) general view of the fracture surface; (b) fibril fracture morphology. The fracture morphology of Specimen 2 is shown in Figure Figure 2.16 (b) shows the overall microstructure, which reveals fibril structures. No crack initiation can be defined in this section of the crack. In the surface near the inner duct wall, fatigue lines can be observed (Figure 2.16(c)). The fatigue lines indicate the existence of fatigue process.

62 47 Figure 2.16 Specimen 2 from sample MB 38-4-A: (a) general view of the fracture surface; (b) fibril fracture morphology; (c) a detailed view of the fatigue line morphology. The fracture morphology of Specimen 3 is shown in Figure The fracture surface also reveals the fibril structure (Figure 2.17 (b)). There is a hemispherical pattern on the fracture surface. The center of the circles is approximately at point A near the inner surface of the duct wall, and this is believed to be the initiation point of the crack. The detailed view of area A shows an impurity (Figure 2.17 (c)).

63 48 Figure 2.17 Specimen 3 from sample MB 38-4-A: (a) general view of the fracture surface; (b) fibril fracture morphology; (c) a detailed view of area A. On the fracture surface of Specimen 4 shown in Figure 2.18, the fibril structure (Figure 2.18 (b)) was also observed. Some fatigue lines were observed at the inner duct surface near area B ; these are shown in Figure 2.18 (c). Figure 2.18 (d) shows an impurity located at both areas A and B.

64 49 Figure 2.18 Specimen 4 from sample MB 38-4-A: (a) general view of the fracture surface; (b) fibril morphology; (c) a close view of fatigue lines near the inner duct surface; (d) a close view of the area A. Sample MB A Sample MB A (Figure 2.19) also shows two cracks, C1 and C2. They occurred at the 10 o clock position. Crack C1 is about 14 inches long on the inner surface and 13 inches on the outer surface Crack C2 is about 40 inches on the inner surface and 38 inches on the outer surface. The two cracks are separated from each other about one inch of distance, measured from the inner surface.

65 50 Figure 2.19 Drawing of sample MB A Four specimens were taken at different positions along the crack in Sample MB A. Specimens 1 and 2 are located near two ends of Crack C1. Specimen 3 is in the middle of Crack C2 and Specimen 4 is close to the end of Crack C2. Figure 2.20(a) shows the general view of Specimen 1. The morphology of Specimen 1 is dominated by fibril structure (Figure 2.20 (b)). Area A appears to be the initiation point, and the close view of area A shows an impurity (Figure 2.20 (c)).

66 51 Figure 2.20 Specimen 1 from sample MB A: (a) general view of the fracture surface; (b) fibril fracture morphology; (c) a detailed view of area A. Figure 2.21 shows the fracture surface of Specimen 2. Impurity and fibril structures in area A can be observed in Figure 2.21 (b) and Figure 2.21(c), respectively. In area C, fatigue lines are revealed (Figure 2.21 (d)). In area B, flaky structure is observed (Figure 2.21 (e)), which indicates an impact failure.

67 52 Figure 2.21 Specimen 2 from sample MB A: (a) general view of the fracture surface; (b) fibril fracture morphology in area A ; (c) a detailed view of the impurity in the area A ; (d) a detailed view of the fatigue line in area B ; (e) the flaky fracture morphology of rapid crack propagation in area C. The morphology of Specimen 3 is shown in Figure The fracture surface is covered by the fibril structure (Figure 2.22 (b)). Also, the fracture surface is covered by many white particles, which make the observation difficult.

68 53 Figure 2.22 Specimen 3 from sample MB A: (a) general view of the fracture surface; (b) fibril fracture morphology; (c) a detailed view of the area A. Figure 2.23 shows the microstructure of Specimen 4. The overall fracture morphology is a fibril structure (Figure 2.23 (b)). An initiation point can be seen in area A, where an impurity is located (Figure 2.23 (c)). Figures 2.23(d) and (e) show the detailed view of areas of B and C, respectively. Both of these areas reveal the existence of fatigue lines.

69 Figure 2.23 Specimen 4 from Sample MB A: (a) general view of the fracture surface; (b) fibril fracture morphology; (c) a detailed view of the imperfection in the area A ; (d) a detailed view of fatigue line in area B ; (e) a detailed view of fatigue line in area C. 54

70 Cracked samples from the SSK Bridge The SSK Bridge also presented severe cracking problems in the ducts in both the column and span parts of the bridge. Sample SSK 131-SB-SE-6 from the column was chosen for presentation. Figure 2.24 shows a sketch of the crack in Sample SSK 131-SB-SE-6, which is about 9 inches long on both inner and outer surfaces. The crack runs longitudinally along the duct. Figure 2.24 Drawing of column sample SSK 131-SB-SE-6 Three specimens were taken from the crack for microstructure examination. Specimens 1 and 2 are located in the middle section of the crack, and Specimen 3 at the left tip of the crack. Figure 2.25 (a) shows the general view of Specimen 1. The crack initiated from the inside surface and propagated through the thickness of the duct wall as suggested by the hemisphere pattern on the fracture surface. Figure 2.25 (b) reveals the fibril structure on the fracture morphology. The imperfection can be found at the initiation point shown in Figure 2.25 (c).

71 56 Figure 2.26 (a) shows the general view of Specimen 2. Figure 2.26 (b) is an imperfection existing in area A. The impurity is thought to be the initiation point for the crack. Figure 2.27 shows the general view of the fracture surface of Specimen 3. Figure 2.27 (b) reveals the fibril structure. Figure 2.25 Specimen 1 from Sample SSK 131-SB-SE-6: (a) general view of the fracture surface; (b) fibril fracture morphology; (c) a close view at crack initiation point.

72 57 Figure 2.26 Specimen 2 from sample SSK 131-SB-SE-6: (a) general view of fracture surface; (b) a close view of the impurity at area A. Figure 2.27 Specimen 3 from sample SSK 131-SB-SE-6: (a) general view of the fracture surface; (b) fibril fracture morphology.

73 Summary of cracking mechanism evaluation The information obtained from the general appearance of the cracks can be summarized as follows: All cracks ran longitudinally along the duct, suggesting the presence of circumferential tensile stress. In ducts taken from the superstructure of the bridge, the majority of the cracking occurred in the region between 10 o clock and 2 o clock, on the top portion of the ducts. Some of the long cracks were formed by the convergence of a series of small cracks. The crack is usually longer on the inner surface than the outer surface, suggesting that cracking was initiated in the inner wall of the ducts. The microstructure of the fracture surfaces provided significant information on the cracking mechanism of the HDPE ducts. The similarities of the fracture morphology of the cracks in both the MB and SSK bridges indicate that the ducts failed by the same mechanism. Fibril structure is the dominant microstructure on the fracture surface, suggesting that slow crack growth is the main mechanism in the cracking. However, impact failure was also involved, as indicated by the flaky morphology. The two long cracks taken from the MB Bridge were formed by the convergence of small cracks, indicating that these cracks were initiated from multiple points.

74 59 The initiation of the cracks was caused by defects such as impurities, air bubbles, or imperfections. These defects were observed at the inner surface of the pipe wall. It appears that cracks mostly likely start from these defects at the inner surface and then propagate through the wall and then in a longitudinal direction. The longitudinal cracking indicates that the driving force is oriented circumferentially. In Samples A and 38-4-A of the MB Bridge, fatigue lines were identified; thus, cyclic loading was also a component of the driving force during the cracking process. It is believed that the cyclic loading is induced mainly by the temperature fluctuations between day and night HDPE duct properties Specification for HDPE materials The current specification used to assess polyethylene pipe materials is ASTM D3350, Standard Specification for Polyethylene Plastic Pipe and Fittings Materials. According to ASTM D3350, polyethylene plastic pipe materials are classified by density, melt index, flexural modulus, tensile strength at yield, environmental stress-crack resistance, and the hydrostatic design basis at 23 o C. The specimens used in all of the tests have the same thermal history, having been taken from mold plaques which were prepared according to ASTM D4703 Procedure C at a cooling rate of 15±5 o C. Each of the properties and the corresponding test method are briefly described in the following section. Density - The density of polyethylene indicates the amount of crystallinity in the material. Higher density indicates larger crystallinity in the material. Either ASTM D1505 or D792 can be used to determine the density. In this study, ASTM D792, Procedure B was used. The test involves immersing the specimen in a liquid which is

75 60 Isopropyl, since the density of HDPE is less than that of water and greater than Isopropyl. The density was obtained by comparing the weight of the specimen in air and in immersed conditions. Melt index (MI) - MI can be qualitatively related to the molecular weight of the polymer. A lower MI value translates to higher molecular weight. The test method for measuring MI is ASTM D1238. The temperature for the test of polyethylene is 190 o C. A load of 2.16 kg (4.76 lb) is applied. The test records the mass that flows out of the die after a period of time. Then the flow rate can be calculated with a unit of g/10min. Flexural modulus - Flexural modulus is the ratio of stress to strain in flexural deformation within the elastic limit. It indicates the ability of the material to resist deformation under load. For polymers, it is closely-related to density; high density translates into high flexural modulus. ASTM D790 uses three-point bending to determine the flexural modulus. Tensile strength at yield The test method is described in ASTM D638. A type IV, dog-bone shaped specimen is applied under tensile load to find the yield point during the process. For polyethylene material, higher density indicates higher tensile strength. Environmental stress-crack resistance (ESCR) ASTM D1693 is the standard for measuring ESCR. In this study, condition C is used: the test specimens are immersed in 100% Igepal at 100 o C. The specimens are inspected periodically for failure.

76 61 Slow crack growth resistance (SCGR) ASTM F1473 (PENT) is the test used by the pressure pipe industry to evaluate SCR. The test involves applying a constant tensile stress of 2.4MPa (348psi) on the specimen at 80 o C in air. Hydrostatic data basis (HDB) ASTM D2837 is the test method used to obtain HDB. The test method involves subjecting the pipe to a series of internal pressures. HDB is the stress that corresponds to a failure time of 100,000 hours. Carbon black content Carbon black is mixed in the pipe resin to increase the UV resistance. ASTM D4218 is used to measure the content of carbon black in the pipe. The specimen is heated in a muffle furnace at 600 o C for three minutes. By doing so, the polymer is burnt off and the residue left is the carbon black. Cell classes are created based on different ranges of values from each of the eight tests. There are six numerical cell classes, and class 0 refers to unspecified. Carbon black content is expressed by a letter. The material properties are specified by a series of cell numbers. Table 2.2 shows the ASTM D3350 designation.

77 62

78 63 Another less-popular specification for identifying the polyethylene material uses types, classes, categories, and grades. This method is described in ASTM D1248, Standard Specification for Polyethylene Plastics Molding and Extrusion Materials. According to this specification, density determines type. The following terms have long been used in practice: Type I Type II Type III = Low Density, when density is between 0.91 and g/cc = Medium Density, when density is between and 0.94 g/cc = High Density, when density is between and g/cc Each of the types is subdivided into three classes according to composition: Class A Class B Class C Natural color Colors including white and black Black, containing not less than 2% carbon black The category is based on flow rate obtained from the melt index test. There are five categories according to ASTM D1248. Category 1 Category 2 Category 3 Category 4 Category 5 flow rate > 25 g/10min flow rate between 10 and 25 g/10min flow rate between 1 and 10 g/10min flow rate between 0.4 and 1 g/10min flow rate > 0.4 g/10min

79 64 Both ASTM D3350 and ASTM D1248 can be used to specify polyethylene materials. ASTM D3350 uses two letters to indicate the material, followed by two numbers indicating density cell and ESCR cell, respectively. Table 2.3 shows the equivalency of the two specifications. Table 2.3 Equivalency of ASTM D3350 and D1248 Specification Grade designations ASTM D3350 PE10 PE20 PE23 PE30 PE33 ASTM D1248 P14 P23 P24 P33 P34 Note: PE = Polyethylene; P = pipe Two more numbers can follow the designation to define the material with greater detail. For example, in PE3408, 08 indicates the hydrostatic design stress of 800 psi Specifications used for the ducts For each of the bridges, material specifications were developed and required for the HDPE ducts prior to the construction. These specifications are showed in Table 2.4. Table 2.5 illustrates the material properties in each of the specifications.

80 65 Table 2.4 Original material specification for the HDPE ducts in each bridge Bridge Source Material Specification Mid Bay Garcon Point Seven Mile Skyway Channel Five Long Key Niles Channel Construction specification Printed on the exterior wall Supplied by duct manufacturer No material specification available Printed on the exterior wall Printed on the exterior wall Printed on the exterior wall ASTM D3350 with a cell classification PE345433C PE3408 ASTM D3350 with a cell classification PE335433C AASHTO specification was used ASTM D3350 with a cell classification PE345433C ASTM 3350 with a cell classification PE335433C PE3406 and PE3408 PE3406

81 66

82 67 All of the above properties except HDB were evaluated in this study. HDB is excluded from the test matrix, since the test must be performed on a complete pipe. In addition to the above tests, the SP-NCTL test (ASTM D 5397) was performed to evaluate the SCR of the duct materials. The specimens were dog-bone. A notch of 20% of the specimen thickness was introduced on one side of the surfaces. The specimens were immersed in 10% Igepal solution at 50 o C, and subjected to 600 psi stress. Five specimens were required for each of the duct samples Test results of the HDPE ducts The properties of retrieved field samples from the seven bridges were tested according to the test methods in the ASTM D3350, together with the SP-NCTL test. The test results were then compared with their corresponding material specification. Table 2.6 shows the test results. All of the field samples conform to the specified values for density, flexural modulus, and tensile strength. However, samples from the MB and SSK Bridges exhibited higher MI values than the specified values. Additionally, all the samples from the MB Bridge failed the ESCR requirement. The SCR of the HDPE ducts from the seven bridges can be assessed by comparing failure times of the SP-NCTL tests. The shortest failure time is obtained in ducts from the MB Bridge, followed by those from the SSK Bridge. Ducts from the Seven Mile, Channel 5 and Nile Channel Bridges show a significantly longer failure time than those from the MB and SSK bridges. Nevertheless, the longest failure time was measured in ducts made

83 68 from pressure-rated resins (PE3408). The SP-NCTL results demonstrate a good correlation with the field performance of the ducts. A poor SCR led to cracking in the field for ducts in both the MB and SSK Bridges, while fewer cracks or none were found in other bridges. Test Method MB Table 2.6 Test results for the duct material properties Garcon Point Seven Mile Result SSK Channel-5 Long-Key Niles Channel Density Passed Passed Passed Passed Passed Passed Passed Melt Index Higher Passed Passed Majority Higher Passed Passed Passed Flexural Modulus Tensile strength Passed Passed Passed Passed Passed Passed Passed Passed Passed Passed Passed Passed Passed Passed ESCR Failed Passed Passed na na na na Carbon black Lower Passed Passed Lower Some lower Some lower Some lower SP-NCTL (hr) at 15% σy 3 to 4 > to 11 na na na SP-NCTL (hr) at na na 13 25% σ y Note: na = not available 16 (PE3406); 243 (PE3408) 50-70

84 Correlation of SCR to other material properties of the HDPE ducts It is known that the SCR of HDPE is in some way related to the molecular weight and crystallinity of the material. The molecular weight of the polymer can be assessed by the MI test and the crystallinity by the density. The SCR of the material is determined by the SP-NCTL test. The correlation between SCR and molecular weight is evaluated by plotting the MI value against the failure time for the SP-NCTL test. The data certainly show a trend indicating that a high MI value tends to yield a short failure time (Figure 2.28). MI (g/10min) Failure time (h) 40 Figure 2.28 Correlations between SCR and MI In the same manner, the correlation between SCR and crystallinity is evaluated by plotting the density value against the failure time for the SP-NCTL test. Figure 2.29 shows the graph plotting density against failure time. The relationship between these two

85 70 parameters is relatively poor, although materials with a lower density tend to yield a longer failure time Density (g/cc) Failure time (h) Figure 2.29 Correlation between SCR and density Although both molecular weight (MI) and crystallinity (density) can affect the SCR, the scatter makes it difficult to predict confidently the SCR of the material Recommended specification As shown in Table 2.5, there is no unified specification for HDPE ducts used in segmental bridges. Based on the previous study, a reliable and unified specification is recommended, which is shown in Table 2.7. Compared with the original specifications, the required density, tensile yield strength, and carbon black content remain the same. Original specifications have MI values in cell 3 or 4. Recommended MI was unified in cell 4 to obtain a relatively high molecular-

86 71 weight material. Original specifications required flexural modulus in cell 4 or 5. The unified specification suggested cell 4, since it is not necessary for the ducts to have high flexural modulus. HDB was raised to cell 4 in the recommended specification for conservative reasons. Furthermore, the current specifications require the ESCR test to assess the SCR property. As stated in Chapter 1, the ESCR test is known to have large standard deviation and is inadequate to distinguish today s HDPE resins. Thus, alternative tests should be implemented in the recommended specification. The SP- NCTL test used in this study has been proven to be able to distinguish the SCR property of different duct materials; however, the SP-NCTL test is not used by the pressure pipe industry; instead the PENT test (ASTM F 1473) is commonly used as the QA/QC test. The correlation between the SP-NCTL test and the PENT test was performed on the duct sample from the Garcon Point Bridge. A failure time greater than 100 hours, which corresponds to cell 6, was obtained from the PENT test. Table 2.7 Recommended specification for HDPE ducts Property Test Method Cell class Required Value Density (g/cc) Melt index (g/10 min) Flexural modulus (psi) Tensile yield strength (psi) Slow crack growth (h) ASTM D 1505 or D 792 ASTM D > < 0.15 ASTM D ,000 - <110,000 ASTM D 638 Type IV ASTM F 1473 (PENT Test) < HDB (psi) ASTM D Carbon black content ASTM D 1603 C > 2%

87 Assessment of stresses in the HDPE ducts The cracking should occur where the largest tensile stresses are located if the material is isotropic. As suggested by the microstructure evaluation in Section 2.4, the driving force is believed to be induced by temperature changes. The distribution of the temperatureinduced stress in the duct was evaluated by both a laboratory-simulated performance test and the Finite Element Method (FEM). In addition, the residual stresses in the duct were measured. As stated in the introductory section, the stress evaluation part of this project was performed by Dr. Hartt s group at Florida Atlantic University (FAU). Their study is summarized in Section and Temperature-induced stresses Laboratory test model A simulated tendon consisting of steel-grout-duct was built by Dr. Hartt s group at FAU to simulate the field tendon (Hartt et al., 2004). The model was 22-inch long with a nominal diameter of 4 inches. The duct was manufactured using pressure-rated resin of PE3408 with a SDR of 21. The diameter of the strands was 0.5 inches. Strain gages and thermocouples were mounted at various locations on the assembled tendon, as shown in Figure Figure 2.31 indicates the locations of the strain gauges mounted on the HDPE ducts. In addition, a void was produced intentionally on the top part of the hydrated grout due the bleeding water during grouting process, as can be seen in Figure 2.32.

88 73 Figure 2.30 Test model built by FAU (From Hartt et al., 2004) Figure 2.31 Strain gage arrangements in the test model (From Hartt et al., 2004)

89 74 Figure 2.32 Void in the test model (From Hartt et al., 2004) The temperature cycles were introduced according to the following four steps: 1. The specimen was grouted at 60 psi at temperatures of 22 to 24 o C, and the pressure was maintained for approximately 22 hours to allow the grout to hydrate. 2. The specimen was then placed in a freezer at -39 C ± 2 C for seven days. 3. After the freezing cycle, the specimen was exposed outdoors for seven days at temperatures between 20 and 35 C. 4. The fourteen-day thermal cycle was repeated to monitor the strain changes. Because the simulated tendon was prepared at 22 to 24 o C and the coefficient of thermal expansion of the duct is much greater than that of the grout, the measurements from strain gages at temperatures higher than 24 o C reflected thermal expansion of the duct only. Conversely, at temperatures below 22 o C, i.e., from -39 to 22 o C, the duct was subjected to

90 75 internal stress from the grout. The actual change in temperature in each thermal cycle was from 61 to 63 o C. Compared to the actual field environment, the maximal range of temperature changes varies from 36 o C (2 to 38 o C) in the Keys region to 50 o C (-9 to 41 o C) in the northern region. Therefore, the laboratory simulation test was performed in a temperature range 20 to 70% higher than service conditions in the field Results from test model Figure 2.33 shows a plot of micro-strain versus time of three gage locations (SG1, SG2 and SG3). The results indicate that SG2 and SG3 have similar tensile strains. Both of them have the highest strain at low temperatures and the lowest at high temperatures. The strain variation range for SG2 is smaller than SG3. SG1, however, has a compressive strain, which could be attributed to the existence of the void below SG1 and SG2 (Figure 2.31). Thus, the induced strain from SG1 and SG2 are thought to be affected by the additional bending effect. For SG3, the strain change is induced solely by the relative contraction of the duct and grout. If the modulus of 1.46x10 5 psi is used for the duct material, the temperature-induced stresses for SG1, SG2, and SG3 are -1020, 440, and 730psi, respectively.

91 76 Figure 2.33 Strain gage measurements (From Hartt et al., 2004) FEM analysis As discussed previously, as temperature decreases, the thermal contraction of the three components in the tendon will not be the same due to their different coefficients of expansion. The coefficient of thermal expansion for HDPE is much larger than those for the steel and grout (the linear coefficients of expansion for grout, steel, and HDPE being approximately 10x10-6 in/in/k, 11.5x10-6 in/in/k, and 117x10-6 in/in/k, respectively.) Because of this, the duct shrinks more than the grout when temperature drops, which induces a tensile hoop stress in the duct. In order to find out the distribution of stresses in the pipe wall, ANSYS, a FEM program, was utilized. A two-dimensional model was established, as shown in Figure The

92 77 model was a two-dimensional representation of the cross-section of a SDR 21 HDPE duct with an interior grout void comprised of 3, 8, or 10% of the interior volume. Triangular, six-node elements were implemented in the analysis. The general illustration of the mesh is shown in Figure Stress-strain response of both grout and HDPE was considered to be linearly elastic. The modules used in the model are 1.46x10 5 psi for the HDPE and 4.06x10 5 psi for the grout. Poisson s ratios for the duct and grout were taken as 0.28 and 0.21, respectively. Figure 2.34 Two-dimensional model of duct-grout system

93 78 Figure 2.35 Mesh in analytical model Results of the analytical model are shown in color-contoured plots of the first principal (hoop) stresses. Figure 2.36 shows the result with 10% void, from which stress distribution can be seen in the color configuration. The tensile stresses were computed at locations SG2 and SG3 on the simulated tendon. The maximum tensile stresses are located at SG2 as well as on the outer duct surface at two sides of the grout void corner. Table 2.8 lists the FEM-computed stress at each of these three locations for three different void volumes. Correspondingly, in Figure 2.37, the stresses are plotted at each of these three locations as a function of void size. For the smallest void size (three percent), stress was greatest at the corner locations followed by SG2. These stresses decrease with increasing void size.

94 79 Figure 2.36 First principal stress distribution when the void is 10% Table 2.8 Summary of the analytical model results Void volume (%) Computed stress (psi) SG2 SG3 Corner

95 80 Principal stress (psi) Corner SG2 SG Void volume (%) Figure 2.37 Void size effects on the principal stress The results of the FEM suggest that the most possible areas that initiate the SC is at SG2 and at the outer duct surface of the void corners. Both the FEM and simulated test indicate that the tensile stresses also distribute around the duct section that contact with the grout Residual stresses It is well-known that residual stresses are generated in pipes during the manufacturing process. The measured residual stresses in HDPE pipes generally fall into the range of psi. The magnitude is certainly too significant to be neglected. Therefore, the effect of residual stress on the material property must be evaluated. The simplest method for measuring the residual stress in the small diameter duct is to slit the duct and measure the diameter change. Dr Hartt s group measured the residual stresses on a section of 10

96 81 inch-long PE3408 duct with a SDR of 21 (Hartt et al., 2004). Figure 2.38 illustrates the splitting test method. Figure 2.38 Measurement of residual stresses by slitting (From Hartt et al., 2004) The change in outside diameter was measured after the splitting. The residual stresses were calculated according to the equation presented in the standard ASTM E Where, Mc Et D1 D0 σ = = ± (2.1) 2 I 1 µ D1 D0 M is the residual moment in the duct, c is the distance from the neutral axis to the point of maximum strain, I is the moment of inertia of the cross-section of the specimen, E is the Flexural Modulus, t is the specimen thickness, µ is Poisson s ratio, D o is the mean outside diameter before splitting, and D 1 is the mean outside diameter after splitting.

97 82 The residual stress measured on the duct was 5.42 MPa (786 psi), with tensile mode on the inner surface and compressive mode on the outer surface. Such stresses are on the same order of magnitude as those obtained from the test model and FEM model. Combining both residual stresses and thermal-induced stresses, it is clear that the highest tensile stress was on the inner surface of the duct. Table 2.9 and Figure 2.39 show the results. This is consistent with the field observation that cracks always originate in the interior surface of the ducts. Table 2.9 Combined temperature-induced stresses and residual stresses Void volume (%) SG2 Corner Stress (psi) Inner duct surface Outer duct surface

98 83 Principal stress (psi) Void volume (%) Corner SG2 Outer duct surface Inner duct surface Figure 2.39 Combined temperature-induced stresses and residual stresses 2.7 Evaluation of fatigue failure Fatigue test Since fatigue is one of the driving forces behind the cracking of the HDPE ducts, the fatigue property of the HDPE materials was evaluated. Duct samples were cut into small pieces, and then compression-molded into 2.5 mil-thick plaques. Specimens 4 inches long by 1 inch wide were cut from the compression-molded plaque prepared by ASTM D4703 procedure C. A 0.5 mil (20% of thickness)-deep notch was introduced at the center of the specimens, at a notching rate of 0.2mil/min. Fatigue tests were performed by three-point bending using an Instron Model 1331 machine under load control. The frequency adopted is 3 Hz, so that no significant heat can be accumulated, and the test can be completed in a short time period as well. To maintain the specimen in position, an initial load of 9.9 lb was applied. The loading was performed by using a half-sine wave loading function. Fatigue failure was defined as the number of cycles to reach a vertical

99 84 deflection of 1 inch. Maximum loads of lb were employed. Figure 2.40 shows the test equipment. Figure 2.41 shows the specimen under testing. Figure 2.40 Fatigue test system Figure 2.41 Specimen in the fatigue test

100 85 The tensile stress σ can be calculated by 3PL σ = (2.2) 2 2bd Where: P = applied load (lb), L = span length (2 inches), b = width of specimen cross-section (1inch), and d = depth of specimen cross-section (0.25 inch) Discussion of fatigue test results Results of the fatigue tests are presented as a plot of stress range versus cycles-to-failure (S-N plot). The relationship conforms to a Power Law equation expressed as b σ = AN (2.3) Where: σ = stress, N = cycles-to-failure, and A and b are material constants, The S-N curves on a log-log coordinate for five duct materials were plotted in Figure The MB and SSK specimens exhibited similar fatigue behavior, having the lowest fatigue resistance. The PE3408 duct specimens exhibited the highest resistance. The specimens from the Seven Mile and Long Key Bridges were intermediate. This ordering of fatigue resistance (i.e. cycles to failures) agrees with the SCR measured by the SP- NCTL test.

101 86 Combining the maximum tensile stresses (about 1200 psi) evaluated from FEM and the measured residual stress (786psi) suggests that the maximum tensile stresses in the duct can be as high as 2000 psi psi is then used as the reference stress to compare the cycles to failure for the different duct samples in the fatigue tests. The plots in Figure 2.42 indicate that a stress level of 2000 psi led to the failure of the SSK and MB specimens after 300 and 900 cycles, respectively. Considering that the in-service ducts experienced one thermally-induced stress cycle per day, the cracking could occur in the ducts in less than one and three years after construction, respectively. Specimens from the Long Key Bridges had cycles to failure of 2.0x10 4 at 2000 psi, which translates to a service life of 50 years in the field. The Seven Mile Bridge sample exhibited 4.0x10 4 cycles. Therefore, 100 years of service life could be possible. For PE3408 resin,, the cycles to failure at 2000psi were 1.0x10 6, making it the most reliable resin for the ducts used in segmental bridges. 1.0E+04 Stress range (psi) 1.0E E E E E E E E E E+07 Cycles to failure Sunshine Skyw ay Seven Mile Mid Bay Long Key PE 3408 Figure 2.42 S-N curves on logarithmic scale for various duct samples

102 Summary This chapter focuses on the SC of the HDPE ducts in segmental bridges. The cracking of the ducts is critical to the safety of the bridges, since the ducts protect the steel strands supporting the superstructure of the bridge. Seven bridges in Florida were involved in this study. The ducts in the MB Bridge and the SSK Bridge presented severe cracking problems. Consequently, the cracking was examined both macroscopically and microscopically. Observation shows that cracking starts on the inner surface of the ducts and propagates through the wall thickness and longitudinally. The microstructure shows that fibril structure dominates the fracture morphology, suggesting the SCG failure mechanism. The fatigue lines can also be found on the fracture surface. It is believed that the driving force for the SC is day-night temperature cycles. To evaluate the effect of duct material properties on SCR, the properties of the materials were tested according to the test methods in ASTM D3350. The results show that some of the properties did not meet the construction specifications. Also, the ESCR test used in these specifications is thought to be inadequate to assess the SCR. A unified specification was then recommended. The major modification in the recommended specification is to replace the ESCR with the SP-NCT test, because it can effectively distinguish the SCR of different ducts, and its results exhibit a good correlation with the field performance. However, the SP-NTCL test is not used by the pressure-pipe industry. Alternatively, the PENT test was adopted in the unified specification. The difference between the SP-NCTL and PENT tests is the test environment. However, they are

103 88 designed based on the same loading mechanism--constant tensile loading. Therefore, either of them is more appropriate to evaluate SCR than the ESCR test. It is of interest to know the stress distribution through the ducts. The laboratorysimulated test and FEM were thus used. Results of both analyses confirm that the most likely initiation positions of the cracking in the duct are the inner surface in the void region. Combined with the residual stresses, large tensile stresses are distributed in the inner surface of the ducts, while the outer duct surface has either compressive stresses or very low residual stresses. These stress analyses agree with the field observation, since the cracking happens where the maximum tensile stresses are distributed. Neither the SP-NCTL nor PENT tests can simulate the cyclic loading occurring in the field. Fatigue tests were thus used to simulate the fatigue process. The cycles to failure were used to estimate the failure times in service conditions. The results show that duct materials from the MB and SSK bridges have the smallest number of failure cycles, and PE3408 has the largest. This presents a good correlation with the field observation, which shows that the MB and SSK bridges have severe cracking problems, while other bridges ducts have minor or no cracking problems. It is concluded that the fatigue test can be used as an effective approach for evaluating the SCR of the HDPE duct material.

104 89 Chapter 3: Stress Cracking in Corrugated HDPE Pipes 3.1 Introduction Corrugated HDPE pipes are the most widely used plastic pipes for drainage applications due to their inherent advantages discussed in Chapter 1. However, like other types of HDPE pipes, the long-term properties of corrugated HDPE pipes are cause for concern. A recent report by Hsuan and McGrath (1999) revealed widespread cracking of corrugated HDPE pipes. Although these pipes are intended to last for approximately 50 years, significant cracking had occurred within a period of a few decades after installation. Unlike the longitudinal cracking observed in the HDPE ducts used in segmental bridges, cracking in drainage pipes was oriented circumferentially within the pipe liner. This pattern suggests that the primary stress acting in the pipes is longitudinal, in contrast to the circumferential stress experienced by HDPE ducts. The purpose of this part of the study was to establish reliable tests and methods to analyze SCR and guarantee a satisfactory lifetime of corrugated HDPE pipes. 3.2 Background Corrugated HDPE pipes Corrugated HDPE pipes are typically used in drainage applications. Most pipes are covered by soil and subjected to substantial compressive loads. Since HDPE is a flexible material, deflection under load is expected. While rigid pipes are considered to be under structural distress when the deformation is above 2% of the pipe diameter, the maximum allowable deformation for flexible HDPE pipes is 5% of the pipe diameter, according to

105 90 AASHTO design standards. Deformation above maximum allowable levels may lead to stability issue of the soil/pipe system, even in the absence of cracking. Corrugated HDPE pipes can range from 4 to 60 inches in diameter, and the liner thickness is relatively thin, regardless of diameter, ranging from approximately 0.06 to 0.15 inches. The pipe stiffness is a function of the corrugation profile, of which there are three types: type C, S, and D. In this study, only type S corrugated pipes were evaluated. In cases where the pipe is buried underground, the pipe and the surrounding soil interact with each other as a composite structure, in which the stiffer element (soil) will respond to a greater fraction of the load. In order to decrease the load transferred onto the pipe, the surrounding soil must be well compacted (95% standard compaction).. Due to the viscoelastic properties of HDPE, the pipe would experience two types of timedependent behavior, creep and stress relaxation, which are shown in Figure 3.1. The load generated at the time of placement is expected to determine the long-term deflection of the pipe and ensure that the pipe does not creep. After the deflection reaches equilibrium, stress relaxation takes place and the stress in the pipe gradually decreases with time.

106 91 (a) (b) Figure 3.1 Creep and stress relaxation behaviors of HDPE pipes Failure in corrugated HDPE pipes According to the report by Hsuan and McGrath (1999), cracked pipes were found in 20 out of 62 sites surveyed in various regions of the US. The majority of the cracks occurred along the circumferential direction on the pipe, indicating the existence of a longitudinal tensile stress. Figure 3.2 shows the cracking in one of the field pipes. Further examination indicates that most of the cracking occurred in the liner, adjacent to the junction, and that the cracks grew from the outer pipe liner surface through the thickness of the liner. Figure 3.2 Cracking in a field corrugated HDPE pipe

107 92 SEM was used to examine the fracture morphology, and revealed that the cracks were dominated by fibril structures, as shown in Figure 3.3. This suggests that SCG was the cracking mechanism. Figure 3.3 Typical fracture morphology from a cracked field corrugated pipe (From Hsuan and McGrath, 1999) Many of the cracking problems were ascribed to improper installation, which results in excessive deflection and buckling of the pipe. However, SC was observed in situation that the deflection of the pipe is less than 5%. Therefore, both the quality of the material and installation are critical in order to achieve a crack-free lifetime. Corrugated HDPE pipes have complex wall geometry, as illustrated in Figure 3.4 depicting a type S corrugated pipe. The complex geometry introduces large variability in residual stresses at different parts of the pipe. Therefore, an adequate performance test should use finished pipes to properly account for pipe geometry and residual stresses.

108 93 Figure 3.4 geometry of type S corrugated pipe 3.3 Test Material The tests were performed on several type S commercial pipes, ranging in diameter from 24 to 60 inches. These pipes included A24, A24, A36, A48, A60; B24, B36, B48, B60, in which the first letter denoted the manufacturer and the subsequent digits indicated the pipe diameter in inches. Most of the tests in this study were conducted on specimens from pipe A36 and A24. The properties that are closely related to SCR are presented, i.e., density, MI, and NCLS. The current specification for the HDPE drainage pipe is AASHTO M 294 specification. M294 requires cell classification of 3 for both density and MI. M294 requires the NCLS test performed on five specimens under an applied stress of 600 psi (4.1 MPa); the average failure time of five test specimens must be greater than 24 hours for virgin pipe resins with no single specimen failure time less than 17 hours.

109 94 Table 3.1 illustrates the tested values. The results indicate that all the pipes meet the density and MI values. However, A60, B24, and B48 pipes do not meet the NCLS requirements. Table 3.1 Properties of the studied pipe samples Sample Density (g/cc) MI (g/10min) NCLS on plaque (hour) A24 A36 A48 A60 B24 B36 B48 B60 Test value Classification 3 3 pass Test value Classification 3 3 pass Test value Classification 3 4 pass Test value Classification 3 4 fail Test value Classification 3 3 fail Test value Classification 3 3 pass Test value Classification 3 3 fail Test value Classification 3 3 pass

110 SCR of liner Introduction The liner test focused on the liner of the pipe. This test was conducted according to the NCLS protocol, with the exception of pre-test preparation of the specimens. Instead of taking the test specimens from the compression molded plaques, the specimens were taken directly from the liner part of the pipe. A notch of 20% of liner thickness is introduced in the center of the specimen. Figure 3.5 shows the specimen location, and Figure 3.6 shows the specimen dimension. Figure 3.5 Liner specimen locations Figure 3.6 Liner specimen

111 96 There are four different configurations in preparing the liner specimen for the NCLS test, which are categorized by the following names: LO specimen cut longitudinally and notched on outer surface; LI specimen cut longitudinally and notched on inner surface; CO specimen cut circumferentially and notched on outer surface; CI specimen cut circumferentially and notched on inner surface. The SCR of the pipe liner is governed by the weakest configuration of the four. Therefore, the configuration that exhibits the least SCR would be used to assess the longterm SCR of the pipe. The average failure times are listed in Table 3.2. The results show that the minimum failure times tended to occur in the LO specimens. Consequently, all liner tests were performed on specimens using type LO configuration for the evaluation of SCR in the involved pipes. Table 3.2 The failure times of four specimen configurations Specimen Failure time tested in Igepal solution at 50 o C (hour) A24 A36 A48 A60 B24 B36 B48 B60 LO LI CO CI Note: -- indicates that the tests were not performed due to the shortage of test material

112 Experimental design for liner test The liner test was developed to evaluate SCR of the finished pipe by incorporating the processing effect into the test. In this study, a number of approaches were adopted to ensure that testing times would be within a practical range. These approaches included increasing test temperatures, using surfactants, and introducing notches into the liners. The following section describes the different test environments. Test agent The use of a 10% Igepal CO630 solution is widely-accepted for SC tests. Although the Igepal solution accelerates the SC process, this does not accurately represent natural conditions. Therefore, extrapolation of experimental data to field performance becomes problematic. A more accurate representation would involve tests that incorporate environmental factors, such as soil, water, and air that are normally in contact with pipes in the field. Since HDPE is highly resistant to chemical corrosion, soil poses little risk to pipe longevity, therefore, only water and air were used in the tests. For comparison, tests in Igepal solution were also performed. Temperature Research had shown that HDPE undergoes material changes in temperatures above 80 o C. Therefore, temperatures were kept below 80 o C for all tests. However, temperatures that are too low risk increasing testing times to levels that are impractical. Tests conducted in water or air requires relatively high temperatures to ensure relatively short failure times. High temperatures are also preferable for testing performed in the presence of Igepal solution; however, temperatures above 50 o C are known to inhibit the effect of Igepal-

113 98 CO630. Therefore, test environments were selected that represent a compromise of effectiveness and practicality. Table 3.3 summarizes the environments used for the liner tests. Table 3.3 Test environments for liner test 10% Igepal Test agent Water Air Solution Temperature ( o C) In each test environment, pipes were subjected to a series of stress levels in order to obtain a full ductile-to-brittle curve. The range of stress levels were determined by the failure times, which were designed to fall within 200 hours in order to make the test practical. Thus, the stress levels used in these tests were between 300 and 1600 psi, at increments of 100 psi. Liner tests were performed on A36 and A24 pipes. For the A36 pipe, the test was performed in all three test agents in order to generate the full ductile-to-brittle curve. For the A24 pipe, the test was only performed in water environment and only the brittle region was generated. The test results for the A36 pipe are presented in three parts. The first part represents the results of tests conducted in water at all three temperatures, the second part represents the results of tests performed in air at all three temperatures, and third part represents data from tests conducted in 10% Igepal solution at 50 o C. The

114 99 results of these tests were compiled in order to compare the effect of the test conditions on the stress cracking Data and Analysis Determination of the ductile-to-brittle curve The test data were assembled in graphical form by plotting log stress against log failure time. The ductile-to-brittle curves were determined using the analytical method described in ISO 9080 Annex B (3). Briefly, this method utilizes a trial-and-error approach to establish the transition point that separates the ductile and brittle portions of the curve, in which 50 points at even increments over the range of log stresses are selected. By assuming that any of these 50 points could be the transition point and assigning the test data in two groups (corresponding to the ductile and brittle behavior, respectively); the data are fitted into the model shown in Equation 3.1: ( log σ ) e log 10 t = c1 + c2 log10 σ + c3 10 log10 σ (3.1) i k + Where: t = time to failure, in hours, σ = applied stress, in psi, σ k = assumed knee stress, in psi (knee stress is the transition stress at which failure changes from ductile to brittle mode) c 1, c 2, c 3i = constants; i = 1 or 2, corresponding to the ductile or brittle region, respectively. e = error variable with zero mean and constant variance

115 100 For each assumed transition point, the constants were computed by the least squares method. The residual variance was calculated based on the fitted lines. After 50 sets of calculations, the best model was the one corresponding to the minimum residual variance. The transition point and fitted lines for ductile and brittle regions were subsequently established from this data. A Matlab code was developed based on this approach to find the best fitting curves (Appendix A) Test results in water environment The liner tests were performed in a deionized water environment at temperatures of 60, 70 and 80 o C. The test data were fitted with ductile and brittle bilinear lines according to ISO 9080 method. Figure 3.7 presents the results of these tests in a single graph. The ductile-to-brittle transition was well-defined in all three curves. The transition stress and time along with the slopes of the ductile and brittle regions at each test temperature are presented in Table 3.4. The slopes in the ductile region increased with temperature, while the slopes in brittle region were less sensitive to temperature. The curves in the brittle region at 70 and 80 o C had similar slopes, while the slope at 60 o C was slightly lower than the other two.

116 101 at 60 o C at 70 o C at 80 o C Figure 3.7 A36 liner test at 60, 70, and 80 o C in water Table 3.4 Summary of fitted curves of liner tests on A36 Environment Temperature ( o C) Stress (psi) Transition Point Time (hr) Ductile Region Slope Brittle Region Water Air % Igepal

117 Test results in air environment Liner tests were also performed in forced air ovens at temperatures of 60, 70 and 80 o C. Figure 3.12, 3.13, and 3.14 show the test data in air and the fitted curves at the three individual temperatures, respectively. Figure 3.8 depicts these results in a single graph. The data suggested that the results at each temperature could be fitted with ductile and brittle bilinear lines. The ductile-to-brittle transition was well-defined in all three curves. Similar to the water data, the slopes in the ductile region increased with temperature. The slopes in the brittle region at 60 o C were significantly lower than those at 70 and 80 o C. at 60 o C at 80 o C at 70 o C Figure 3.8 A36 liner test at 60, 70, and 80 o C in air

118 Test results in Igepal solution The test data and the fitted curve for the tests performed in Igepal solution at 50 o C are presented in Figure 3.9. The transition stress, time, and slopes in the ductile and brittle regions are shown in Table 3.4. Figure 3.9 A36 liner test at 50 o C in Igepal solution Comparing SCR in Different Test Environments The ductile-to-brittle curves in the three different test environments are shown in Figure Table 3.4 summarizes the transition points and the slopes for each test. At each temperature, the ductile-to-brittle curves for tests performed in water and air appeared to

119 104 be similar. The transition stresses and failure times were similar at 60 and 70 o C. In contrast, tests at 80 o C, suggest that the tests performed in water had a transition point at lower stress point and at longer time. The difference between the tests in water and air was reflected in the slopes of the curves. In the ductile region, the curve in air was steeper than that in water, whereas the opposite phenomenon was observed during the brittle region. The curves for the test performed in Igepal solution were significantly different from those in the other two test environments. These curves had a significantly steeper slope than those in air or water. These results indicate that Igepal can significantly accelerate the SC process in HDPE pipes. Tested in Igepal solution at 50 o C Solid line --tested in water or Igepal solution Dotted line--tested in air at 80 o C at 60 o C at 70 o C Figure 3.10 Compiled graph for A36 liner test in all environments

120 105 The difference between the test results under each test condition could also be determined by calculating the corresponding activation energy, which is defined as the minimum energy required to initiate the reaction. The Arrhenius model expressed in Equation 3.2 was used to calculate the activation energies. 1 t f = Ae E RT (3.2) Where: t f = failure time at a specific applied stress, in hr, E = activation energy, in kj/mol, T = test temperature, in K, R = gas constant (8.314 J/mol-K), A = material constant Activation energies are usually determined experimentally by measuring the reaction rate 1 1 ( ) at different temperatures (T), plotting the logarithm of against 1/T on a graph, t f t f and determining the slope of the straight line that best fits the points. Figure 3.11 shows 1 the relation between natural log ( ) and 1/T for the test data in water at 600 psi. Based t f upon the calculated slope of , Equation 3.2 yeilds activation energy of 92.5 kj/mol.

121 ln(1/t) y = x /T Figure 3.11 Activation energy calculation for A36 liner test in water at 600 psi For tests conducted in water, the activation energies ranged from 228 to 232 kj/mol for the ductile curves, and 92 to 104 kj/mol for the brittle curves. For tests conducted in air, the activation energies ranged from 185 to 189 kj/mol for the ductile region, and 100 to 113 kj/mol for the brittle region. A low applied stress yields a high activation energy, and vise versa. Extensive investigations on all polyethylene by Lu and Brown (1986, 1987, and 1990) and Huang and Brown (1988) demonstrated that the activation energy for SCG (brittle region) was approximately 100 kj/mol. Lu and Brown (1990) stated that the activation energy in the ductile region ranged from 172 to 270 kj/mol. Our results strongly recapitulate these previous findings. However, a previous report by Lu and Brown (1990) suggested that the activation energy is independent of stress, while an inverse proportion between the stress and activation energy was found in this study.

122 Residual stress Residual stress measurement The difference in the failure times among the four configurations of liner specimens indicated that the residual stress in the pipe may have an influence on SCR. Thus, the residual stresses in the pipe were measured. However, the geometry and large diameter of the pipe limited the application of the simple slitting method that was used to measure residual stresses for small diameter smooth pipes. As an alternative, residual stresses were measured on specific regions cut from the larger diameter pipes. Although the most popular and accurate method is layer removal, the thinness of the pipe liner (mostly less than 0.12 inches) made this method impractical. Therefore, a thermal annealing method was performed in this study. Unfortunately, this process could potentially result in changes to the molecular structure of the material. However, since the SC tests were performed at relatively high temperatures of up to 80 o C with little change to the molecular structure, annealing below 80 o C is presumed to be acceptable. The annealing approach involves two steps: releasing the residual stress by heating at a certain temperature, and estimating the residual stress by comparing the specimen geometries before and after annealing. Specimens of 3 inches by 0.5 inches were cut from pipe liner in both longitudinal and circumferential directions. If the crest of the corrugation was wide enough, specimens were also cut from the crest. The specimens were subsequently placed in heated ovens for various time intervals. Four trial temperatures were adopted for the annealing: 50 o C, 60 o C, 70 o C, and 80 o C. These temperatures were chosen because they were the temperatures at which the stress

123 108 cracking tests were performed. The annealing interval was determined by direct observation of specimen shape changes. After approximately 6 hours, the specimen shape experienced no further changes at any of the above four temperatures, suggesting that 6 hours of annealing was sufficient to release all residual stress. The initial shape of the specimens was roughly flat, however following the release of residual stress, a curvature on the specimen was observed. The change in the curvature corresponded to a certain amount of strain, from which residual stress could be calculated. The change of the curvature was measured by the dial gauge. Figure 3.12 shows the measurement of the arc height. The procedures that were used to calculate residual stresses are included in Appendix B. Figure 3.12 Measurement of arc height of specimen

124 Residual stresses effect Effect shown from four specimen configurations The residual stresses measured by the annealing method are listed in Table 3.5. All the values shown are residual stresses on the outer surface of the liner. The data indicate tensile residual stresses in the longitudinal direction of the pipes. In the circumferential direction, the stress mode varies from pipe to pipe. Generally, the absolute values of residual stresses in longitudinal direction are much larger than those in circumferential direction. This indicated that the extrusion rate has a larger effect than the cooling process on generating the residual stresses.

125 110

126 111 McGrath (2003) conducted a parametric study using corrugated HDPE pipes in which effects of soil compaction conditions, depth of fill, and support under the pipe haunches was investigated. The result showed that both circumferential and longitudinal long-term service strain should be less than 1.6%, corresponding to a stress of approximately of 320 psi assuming a long-term modulus of 20,000 psi. The measured residual stresses are certainly too large to neglect, especially in the longitudinal direction, where the residual stresses typically range from 40 to 815 psi. The effect of residual stresses can be seen by the NCLS test data included in Table 3.5. These data indicated that the large tensile residual stresses correspond to shorter failure times. Since the largest tensile residual stresses are found on the outer surface of the longitudinally oriented specimens, the shortest failure times always occurred in the LO specimens. However, the longest failure times occurred in the LI specimens, due to the maximum compressive residual stresses on the inner surface. The residual stresses in the circumferential direction are not as significant as those in longitudinal direction. Therefore, the failure times are between those of LO and LI specimens. It should be pointed out that the measured residual stresses are only on the pipe surfaces and that the distribution of the residual stresses across the pipe liner thickness is unknown. Thus, the relationship between the failure times and residual stresses can only be evaluated qualitatively Effect shown from comparison of liner and plaque specimen It is believed that compression molded plaques could relieve most of the residual stresses in the pipe. By comparing the failure times between specimens taken from plaque and liner, the effect of residual stresses on SCR can be assessed. The A36 pipe was used in

127 112 this evaluation study. Table 3.6 presents the parameters for the plaque and liner test results. Figure 3.13, 3.14, and 3.15 compare the test results from both liner test and plaque test results in Igeapl solution, water, and air environments, respectively. The figures show that the two sets of tests have similar transition points at the same test conditions. However, there is a significant difference in the calculated slopes of the fitted curves. The difference in slopes is smaller at 80 o C than at 60 or 70 o C. In the ductile region, plaque tests tended to have steeper slopes, while in the brittle region, plaque tests tended to have shallower slopes. The steeper slopes in brittle region for liner test translated into lower activation energy. The calculated activation energies in the brittle region are listed in Table 3.7. Since the tests were performed under identical conditions, the difference in activation energies between liner and plaque tests are primarily ascribed to the residual stresses in the specimens. Residual stress varies within the specimens, therefore the effect of the residual stresses on cracking initiation and propagation are unclear.

128 113

129 114 Solid line--liner test Dotted line--plaque test Figure 3.13 Comparison of test results from liner test and plaque test in Igepal at 60 o C Solid line--liner test Dotted line--plaque test at 80 o C at 70 o C Figure 3.14 Comparison of test results from liner test and plaque test in water

130 115 Solid line--liner test Dotted line--plaque test at 80 o C at 60 o C at 70 o C Figure 3.15 Comparison of test results from liner test and plaque test in air Table 3.7 Activation energies from different tests Test Test environment Activation energy range (kj/mol) Liner test on A36 Plaque test on A36 water air water air

131 SCR of junction Introduction Although the liner tests offer consistent results with short failure times, the application of these data to the prediction of pipe lifespan suffers from two main limitations. First, these tests do not consider the complexity of the pipe geometry. Field observations indicate that the SC in corrugated HDPE pipes commonly occurs at the junction area between the liner and the corrugation. Second, liner test required a 20% deep notch on the specimen, which is rarely observed in field pipes. Furthermore, the correlation between the notch and the defect in the service pipe remains unclear. The development time for the defect on the pipe surface to 20% thickness represent a significant proportion of total failure time. Junction tests were designed to overcome the shortcomings of the liner test. The discontinuity of the junction generates a stress concentration, making it more susceptible to SC. Figure 3.16 illustrates the configuration of the junction. Figure 3.16 Junction configuration

132 Experimental design for junction test To evaluate the SCR of the junction, specimens were taken from this part of the A24 pipe. Figure 3.17 shows the location of the specimens. Figure 3.18 shows the typical junction specimen. The junction tests were limited to water environment at temperatures of 60, 70, and 80 o C. The applied stresses ranged from 350 psi to 1150 psi with increments of 100 psi. Figure 3.17 Junction test specimen location Figure 3.18 Typical configuration of junction specimen