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SPECIAL SECTION PAPERS: Pipeline Systems

The Effect of Internal Pressure on the Tensile Strain Capacity of X52 Pipelines With Circumferential Flaws

[+] Author and Article Information
Diana Abdulhameed

Department of Civil and
Environmental Engineering,
University of Alberta,
Edmonton, AB T6G 2R3, Canada
e-mail: dabdulha@ualberta.ca

Celal Cakiroglu

Department of Civil
and Environmental Engineering,
University of Alberta,
Edmonton, AB T6G 2R3, Canada
e-mail: Cakiroglu@ualberta.ca

Meng Lin

Department of Civil
and Environmental Engineering,
University of Alberta,
Edmonton, AB T6G 2R3, Canada
e-mail: lin4@ualberta.ca

Roger Cheng

Department of Civil and Environmental
Engineering,
University of Alberta,
7-306 Donadeo Innovation Centre for Engineering,
Edmonton, AB T6G 1H9, Canada
e-mail: Roger.cheng@ualberta.ca

John Nychka

Department of Chemical and Materials
Engineering,
University of Alberta,
9107-116st North West,
12-215 Donadeo Innovation Centre for Engineering,
Edmonton, AB T6G 1H9, Canada
e-mail: jnychka@ualberta.ca

Millan Sen

Enbridge Pipelines Inc.,
10201 Jasper Avenue North West,
Edmonton, AB T5J 2J9, Canada
e-mail: Millan.sen@enbridge.com

Samer Adeeb

Department of Civil and Environmental
Engineering,
University of Alberta,
7-245 Donadeo Innovation Centre for Engineering,
Edmonton, AB T6G 1H9, Canada
e-mail: adeeb@ualberta.ca

Contributed by the Pressure Vessel and Piping Division of ASME for publication in the JOURNAL OF PRESSURE VESSEL TECHNOLOGY. Manuscript received March 20, 2015; final manuscript received April 12, 2016; published online July 22, 2016. Assoc. Editor: Hardayal S. Mehta.

J. Pressure Vessel Technol 138(6), 061701 (Jul 22, 2016) (18 pages) Paper No: PVT-15-1045; doi: 10.1115/1.4033436 History: Received March 20, 2015; Revised April 12, 2016

Wide plate testing has been traditionally applied to evaluate the tensile strain capacity (TSC) of pipelines with girth weld flaws. These wide plate tests cannot incorporate the effect of internal pressure, however, numerical analysis in recent studies showed that the TSC is affected by the level of internal pressure inside the pipeline (Wang et al. 2007, "Strain Based Design of High Strength Pipelines," 17th International Offshore and Polar Engineering Conference (ISOPE), Lisbon, Portugal, Vol. 4, pp. 3186–3193). Moreover, most of the past studies focused on the effect of circumferential flaws on the TSC for pipelines of steel grade X65 or higher. The current Oil and Gas Pipeline System Code CSA Z662-11 provides equations to predict the TSC as a function of geometry and material properties of the pipelines. These equations were based on extensive studies on pipes having grades X65 or higher without considering the effect of internal pressure. This paper investigates the TSC for pipelines obtained using an experimental technique considering the effect of internal pressure and flaw size. Eight full-scale tests of X52 NPS 12 in. pipes with 6.91 mm wall thickness were conducted in order to investigate the effect of circumferential flaws close to a girth weld on the TSC for vintage pipelines subjected to eccentric tensile forces and internal pressure. The tensile strains along the pipe length and on the outer circumference of the pipe were measured using biaxial strain gauges and a digital image correlation (DIC) system. Postfailure macrofractography analysis was used to confirm the original size of the machined flaw and to identify areas of plastic deformation and brittle/ductile fracture surfaces. From the experimental and numerical results, the effect of internal pressure and flaw size on the TSC and the crack mouth opening displacement (CMOD) at failure were investigated and presented.

Copyright © 2016 by ASME
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References

Figures

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Fig. 1

Machined flaw close to the girth weld

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Fig. 2

Base connection showing the pin–yoke assembly connected to the tongue

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Fig. 3

Speckle pattern for the DIC system

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Fig. 4

Representative torch-cut specimens of tested pipe for failure analysis (sample 2 overlaid on sample 1; pen shown for scale)

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Fig. 9

Tensile strain along the pipe length aligned with the flaw (the circle shows the intersection of the average strain with the strain profile)

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Fig. 8

Comparison between the DIC and strain gauge readings at quarter of the length of the pipe. Top refers to the top quarter, and bottom refers to the bottom quarter.

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Fig. 7

Average longitudinal and circumferential true stress–strain curve of the X52 base metal

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Fig. 6

Geometry (1/4 of the pipe + ½ end plate) and mesh of the finite-element analysis model

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Fig. 5

Specimen sectioning progression for failure analysis (sample 2; LHS: left-hand side; RHS: right-hand side). (a) Markings for sections (lines surrounding “XXX” markings indicate rough cuts for hacksaw), (b) after sectioning for fractographic examination, and (c) final sectioning for mounting. Final cuts on surfaces to be mounted and polished for optical metallography were made by a slow-speed diamond saw blade to minimize subsurface damage. The small pieces with “XXX” markings were mounted in epoxy for metallographic examination. Scale in (c) is in millimeter.

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Fig. 10

The development of ε0.5L with the applied load showing the effect of the internal pressure. Each curve shows the variation of ε0.5L up to failure for two tests sharing the same flaw size but different internal pressures.

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Fig. 19

Representative macrophotographs of fractured specimens. (a) Sample 1 fracture surface illustrating region of machined flaw, and flat, brittle final fracture surface decorated with corrosion products from prolonged contact with testing fluid. (b) Sample 2 fracture surfaces; general flat nature of the final fracture surfaces can be observed, and ligaments indicated by white arrows. Note that if the fracture surfaces were to be matched together, the ligaments would form a continuous strip, which indicates that the flaw machining induced multiple cracks at the initiation stage, but then they collapsed into one crack leading to final failure leaving behind the remnant of material between the initial cracks as a ligament.

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Fig. 20

(a) Representative composite optical micrographs of macro-etched metallographic specimens and (b) schematic of manufactured flaw location with respect to the HAZ after testing (samples 1 and 2). Note the large plastic deformation zone beneath the machine flaw notch root where the final fracture surface was created via cracking. In all the eight samples, the machined flaws were not located within the HAZ. The irregularity of the notch shape was due to the flaw machining process creating multiple notches with variable depths.

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Fig. 21

Macro-etched specimen showing representative locations for measurements of features found in Table 4

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Fig. 22

Bar graph of notch depth (as a fraction of original wall thickness) versus sample number. By in large the target values of 25% and 50% were achieved within a few percentage, except for sample 8, which was ∼10% lower than targeted.

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Fig. 11

The development of ε0.5L with the applied load showing the effect of flaw depth. Each curve shows the variation of ε0.5L up to failure for two tests sharing the same flaw length and internal pressure.

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Fig. 12

The development of ε0.5L with the applied load showing the effect of flaw length. Each curve shows the variation of ε0.5L up to failure for two tests sharing the same flaw depth and internal pressure.

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Fig. 13

Variation of CMOD with the applied load. All the curves exhibited the same trends with the same critical CMOD for each flaw size. The number in brackets indicates the distance between the two points used for calculating the CMOD.

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Fig. 14

Comparison between the experimental (solid) and finite-element model (dashed) load–displacement curves for all the tests. Each curve shows two tests sharing the same flaw size but different internal pressures.

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Fig. 15

Comparison between the experimental (solid) and finite-element model (dashed) load–rotation curves for all the tests. Each curve shows two tests sharing the same flaw size but different internal pressures.

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Fig. 16

Comparison between the experimental and finite-element model load-axial strain at quarter the pipe length at angles 90 deg and 270 deg from the flaw location for all the tests

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Fig. 17

Comparison between the experimental and finite-element model axial strain distribution at the onset of failure on the tension side for all the tests

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Fig. 18

Representative composite macrophotographs of flaws after failure under oblique lighting: (a) sample 1 and (b) sample 2. White surfaces with black speckled pattern (top images in (a) and (b)) illustrate little surface deformation on the outside surface, except some at the flaw tips; whereas significant surface deformation (depressions at the flaw tips and walls) is visible on the inside surfaces for both samples. (b) also shows that there were multiple notches present and that some crack mouth opening occurred in more than one effective flaw in some fraction of the flaw length (e.g., left-hand flaw tip in top image in (b)).

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Fig. 23

Optical micrograph of sample 1 at the notch root (cross section). The elongated grains visible in the metal (in the center of the image) are indicative of severe plastic deformation prior to fracture, and such strain was localized beneath the notch root (plastic hinging).

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Fig. 24

Comparison between the experimental and predicted TSC

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