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Fracture Control for the Oman India PipelineT.V. Bruno, Metallurgial Consultants, Inc.
AbstractThis paper describes the evaluation of the resistance to fracture
nitiation and propagation for the high-strength, heavy-wall pipeequired for the Oman India Pipeline (OIP). It discusses the
unique aspects of this pipeline and their influence on fracture
ontrol, reviews conventional fracture control design methods,
heir limitations with regard to the pipe in question, the extent to
which they can be utilized for this project, and other approaches
being explored. Test pipe of the size and grade required for the
OIP show fracture toughness well in excess of the minimum
equirements.
ntroductionThe Oman India Pipeline (OIP) will transport natural gas
pproximately 1100 km from Oman to India under the ArabianSea, at water depths to 3525 meters. Because of the unprecedented
water depth the design requires line pipe of a size and grade never
before manufactured, much less utilized for an offshore pipeline.
The pipe will be API Specification 5L Grade X70, with an inside
diameter of 610 mm and a wall thickness ranging from 36 to
44mm. The maximum hoop stress will be 330.4 MPa (under shut-
n conditions) and the design temperatures are 0C minimum, 50C
maximum.
The pipeline will be constructed with U-O-E pipe made
rom low-carbon, low-sulfur, microalloyed steel plate
manufactured with thermo-mechanical process control (TMPC)
ncluding accelerated cooling. The specified mechanical properties
re shown in Table 1. Because so much of the pipeline will be in
deep water, the hoop stress of approximately 70 percent of the
ength of the pipeline will be less than 50 percent of the specifiedminimum yield strength (SMYS). Therefore, for most of the
pipeline the pot ential for fracture will be much lower than for mostpipelines. Figure 1 shows the maximum hoop stress vs. location
long the pipeline.
Principles of Fracture Control Design
Fracture control design of pipelines requires that under the most
dverse conditions: 1) the pipe has sufficient fracture toughness to
olerate small flaws without fracturing; 2) if the pipe ruptures from
ny cause, the fracture is ductile; 3) the steel has the capacity to
bsorb sufficient energy to arrest a ductile fracture, or crack
rrestors are added.
Considerable research on the behavior of pipelinesponsored by the Pipeline Research Committee of the American
Gas Association, (1) British Gas, (2) the European Pipeline
Research Group (3) and others has resulted in analytical and test
methods to evaluate these three requirements based on the
properties of the pipe and the design of the pipeline. Evaluation of
hese methods by full-scale burst tests as well as their widespread
uccessful application has shown them to be adequate within
ertain limits of operating conditions and pipeline designs.
However, as will be discussed, some aspects of the OIP, especially
the wall thickness and design pressure are outside these lim
Nevertheless, as will be shown, the methods can be conservatapplied to evaluate resistance to fracture initiation and to gi
reasonable estimate of resistance to fracture propagation.
Fracture Initiation.AGA-Battelle Equations. The resistance to the initi
of ductile fractures can be evaluated for through-wall or pa
wall flaws using Equations (A-1) and (A-2) shown in
Appendix, which were developed by Battelle under
sponsorship. These equations give the size of a critical flaw
one that will cause a leak or rupture, as a function of the Charp
notch (CVN) toughness, the pipe size and grade, and the hstress. Similar equations have been developed for high-tough
the pipe size and grade, and the hoop stress. Similar equa
have been developed for high-toughness pipe for which fracinitiation is independent of the CVN toughness but Equation
1) and (A -2) were used because the results are conservative.
As a first approach, critical flaw sizes for the OIP
calculated assuming a CVN fracture toughness of 100
specified for the longitudinal weld seam, as opposed to 200
the base metal, for conservatism. For convenience only thr
wall (T.W.) and 50-percent wall surface flaws are considered.
pipeline has been divided into 17 increments by wall thicknes
design purposes. As shown in Table 2 and Figure 2, the calcucritical flaw sizes are very large, ranging from 254 mm to m
than 1000 mm.
Equations (A-1) and (A-2) have been ver
experimentally only for wall thickness up to 21.9 mm for theand using the hoop stress based on the actual design pressur
can calculate critical flaw sizes within the wall thickness limi
which the equations have been verified experimentally. T
values are very conservative because the assumed wall thick
gives a higher hoop stress than the actual hoop stress.
Table 3 and Figure 3 show the calculated hoop stre
and critical flaw sizes based on a constant wall thickness of
mm. First consider the pipe from KP segments 3 through 15
flaw lengths over this portion of the pipeline are order
magnitude above the limits of detectability by ordinary inspe
methods. Moreover, the assumed wall thicknesses are 39.2 pe
(36.0 to 21.9 mm) to 50.2 percent (44.0 to 21.9 mm) less than
specified wall thicknesses and the hoop stresses are 1.4 totimes the actual maximum design stresses.
Next consider the pipe in KP segments 1, 2, 16, an
Even in these shallow-water areas the flaw sizes assuming a
mm wall thickness are relatively large and well within the limi
detectability. For these segments the wall thicknesses are
percent (38.8 to 21.9 mm) to 46.7 percent (41.1 to 21.9 mm)
than the specified wall thicknesses and the hoop stresses are
1.8 times the design stresses.
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From the above it can be seen that even with conservative
ssumptions the OIP has adequate resistance to fracture initiation,
based on the AGA -Battelle equations.BSI PD 6493. Resistance to fracture initiation can also be
valuated using crack tip opening displacement (CTOD) and the
method of British Standard Institute's PD 6493 : 1991, "Guidanceon methods for assessing the acceptability of flaws in fusion
welded structures"(4) This method is commonly applied to welds
but is equally applicable to the pipe base metal.
Two cases were analyzed, a shallow-water case and a
deep-water case, with the conditions shown in Table 4. The critical
law size was determined for the weld and base metal, and for
nternal and external surface flaws. The results were plotted as
ritical flaw length vs. depth (d) expressed as a fraction of the wall
hickness (t), i.e., d/t, for CTOD values of 0.38 mm and 0.64 mm.
Figure 4 shows the results for the shallow-water weld metal. (The
usps in the curves are due to the formulas for calculating stress
ntensity; in reality the curves would be smooth.) As shown,nternal flaws have a smaller critical flaw size than external flaws
nd are therefore more significant. For the lower CTOD value, the
ritical internal flaw length for deep flaws (>d/t = 0.40) is in theneighborhood of 20 mm and increases rapidly for shallower flaws.
Figure 5 shows the results for the deep-water weld metal, internal
law (the external flaw size, which is larger, is not shown). For
deep flaws at the lower CTOD value, the critical flaw length is in
xcess of 30 mm.
The shallow-water base metal internal flaw case is shown
n Figure 6. At the lower CTOD value, the minimum critical flaw
ength is about 40 mm. The deep-water base metal case gives even
arger flaws and is not shown.The critical flaw sizes for the weld metal are smaller than
hose for the base metal because PD 6493 assumes residual
welding stresses for the former. Also, for the same design
onditions, PD 6493 gives smaller flaw sizes than the AGA-Battelle equations because of more conservative assumptions.
Consequently, the flaw sizes derived from the PD 6493 method
an be considered a lower bound.
Fracture Propagation.The resistance to the propagation of ductile fractures can
be evaluated by comparing the fracture speed to the decompression
behavior of the gas in the pipeline. When a pipeline ruptures, gas
decompression waves at different pressure levels propagate along
he pipeline away from the opening in each direction. Under some
onditions the fracture speed is slow enough that the
decompression wave at the pressure necessary to support fracture
passes the crack tip and the fracture arrests. Under other conditionshe fracture speed is fast enough for the crack tip to always lead the
decompression wave of the pressure necessary to cause arrest and
he crack continues to propagate.
AGA-Battlle Equations. The velocities of gas
decompression and fracture propagation can be calculated using
Equations (A-3), (A-4), and (A-5) in the Appendix, which were
lso developed by Battelle for the AGA. The same data can be
generated using two computer programs, GASDECOM and
DUCTOUGH, available from the AGA.(1)
The programs plot
fracture velocity vs. pressure and gas decompression velocity
pressure on the same curve. For a given pipe size and grade
given operating pressure, the fracture velocity varies inversely
CVN upper shelf toughness. The fracture velocity curve has
shape and levels off at a constant pressure that represent
fracture arrest pressure. The decompression curve is a functiothe gas composition.
When the CVN toughness is such that the two curves are tan
fracture is unstable and will eventually arrest. When the curve
separated, the pressure quickly reduces to the arrest pressure
the fracture arrests quickly. When the curves intersect, the cra
remains at a pressure sufficient to support fracture and propag
continues.
Curves were generated for a shallow-water and a
water case to determine the CVN toughness necessary to prec
long fractures. As shown in Figures 7 and 8, the required u
shelf energies for fracture arrest are:
Shallow-water Case: ~45 JDeep-water Case: ~3.4 J
The required toughness for fracture arrest is extremely low fo
deep-water case and lower than might be expected for the shawater case. One reason for the low values for the deep-water
is the low hoop stress; because of water pressure the tensile
stress is only 22 percent of SMYS.
Both cases are influenced by the fact that the
composition and high pressure are such that the gas is very d
and tends to behave more like a liquid than a gas,
decompression waves travel faster than less dense gases.Crack Tip Opening Angel. The CVN test currently i
most widely used test to evaluate the resistance of pipelinepropagating ductile fractures. Recently a new approach util
the crack tip opening angle (CTOA) has been proposed. (5-7)
this approach, the fracture resistance of the pipe, termed (CTO
is compared to the driving force of the pressurized gas, ter(CTOA) max for a given pipeline design. The equilib
condition for ductile fracture propagation/arrest is:
(CTOA)c= (CTOA)max .(1)
and the condition to preclude propagation is:
(CTOA)c> (CTOA)max (2)
The value of (CTOA)c is determined by dynamic fra
tests using three-point bending specimens of two different liga
lengths and the value of (CTOA)max is determined usi
computer program called PFRAC.
(7)
Ten CTOA tests were run on samples of 660-mm O
41.3-mm wall test pipe with a yield strength of 478 MPa.
pipe had been produced from plate with similar chem
composition and processing as specified for the OIP.
(CTOA)max was determined based on the OIP design conditi
The average CTOA of the ten specimens was 11.7 compare
the calculated (CTOA)max of 3.3. The fact that (CTOA)c was
than three times (CTOA)max indicates that fracture propagati
highly unlikely.
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because it is virtually impossible to assure resistance to fracture
nitiation from all causes, such as marine accidents and other low-
probability occurrences, fracture propagation must also be
onsidered. For the OIP, consideration of fracture propagation is
econdary to consideration of fracture initiation and is an issue
only in the shallow-water areas of the pipeline. Moreover, fracturepropagation is principally an economic consideration relating to
he cost of repairing "long" failures as compared to "short" failures.
Analyses using conventional methods that have not been
verified for the OIP conditions indicate a high probability that the
pipe will have adequate resistance to fracture propagation.
Verification will require an expense that may not be justified and
other means of limiting fracture propagation, such as the use of
rack arrestors may be more practical.
Conclusions .
. Based on conventional fracture control technology using
onservative assumptions, pipe produced to the OIP specificationwill have adequate resistance to fracture initiation under the most
dverse operating conditions.
. Resistance to fracture propagation evaluated byonventional methods is high, however, these methods have not
been verified for the OIP pipe size and grade and operating
pressure. Considering the costs of verifying the resistance to
racture propagation by full-scale testing, the use of crack arrestors
may be more cost effective.
. Tests on a trial production of one kilometer of pipe
howed fracture toughness well in excess of the minimum
equirements of the project.
Acknowledgements
We thank Europipe for conducting the West Jefferson tests.
References
1. Eiber, R.J., Bubenik, T. A., and Maxey, W.A., "Fracture
Control Technology for Natural Gas Pipelines," AGA,
Project PR-3-9113, NG-18 Report No. 208, Dec. 1993.
2. Fearnehough, G.D., "Crack Propagation in Pipelines,"
The Institution of Gas Engineers, March 26-27,1974.
3. Vogt, G.H., Bramante, M., Iones, D.G., Koch, F.O.,
Koglar, J., Pro, H., and Re, G., "EPRG Report on
Toughness for Crack Arrest in Gas Transmission
Pipelines," 3R Internatiof1al (1983) 22 ,98.
4. PD 6493, "Guidance on Methods for Assessing the
Acceptability of Flaws in Fusion-Welded Structures,"BSI, Bulletin Box No. 15A, 1991.
5. Kanninen, M.F. and Grant, T.S., "The Development and
Validation of a Theoretical Ductile Fracture Model,"
Eighth Symposium on Line Pipe Research, AGA -Pipeline
Research Committee, Sept. 26-29,1993.
6. Demofonti, G., Kanninen, M.F., and Venzi, S., "Analysis
of Ductile Fracture Propagation in High-Pressure
Pipelines: A Review of Present-Day Prediction Theories,"
Eighth Symposium on Line Pipe Research, AGA -Pip
Research Committee, Sept. 26-29,1993.
7. Basically, G., Demofonti, G., Kanninen, M.F., and V
S., "Step by Step Procedure for the Two Specimen C
Test,"Pipeline Technology, II, 503.
8. Preston, R., "Improvement in UOE Pipe ColResistance by Thermal Aging," paper OTC
presented at the 1996 Offshore Technology Confere
Houston, Texas, May 6-9.
9. Bruno, T.V., "The Effect of Water Overburden on Du
Fractures in Gas Pipelines," Doc. No. 9100-ALA-R
1001, 1995.
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Yield Strength, MPa
Tensile Strength, MPa
Hardness, HV 10
CVN at -10 deg. C
Energy, J: Base Metal
Weld
% Shear: Base Metal
DWTT at -10 deg. C.
% Shear
CTOD at -10 deg. C, mm
(Weld Metal)
* Avg. of 3/Any 1
TABLE 1 - SPECIFIED MECHANICAL PROPERTIES
100/75 min. *
90/75 min. *
85 min.
0.40 min.
482 min., 586 max.
565 min., 793 max.
248 max.
200/150 min.*
Increment KP % SMYS MPa T.W. d/t = 0.51 0-29 68.5 330.6 254.0 355.6
2 29-42 62.9 303.6 292.1 431.8
3 42-56 60.5 292.0 292.1 495.3
4 56-68 40.2 194.0 457.2 >1000
5 68-278 31.6 152.5 558.8 >1000
6 278-282 21.9 105.7 736.6 >1000
7 282-535 22.4 108.1 736.6 >1000
8 535-611 27.8 134.2 673.1 >1000
9 611-617 27.7 133.7 533.4 >1000
10 617-742 33.1 159.8 558.8 >1000
11 742-755 32.4 156.4 584.2 >100012 755-788 41.8 201.7 431.8 >1000
13 788-854 46.2 223.0 381.0 812.8
14 854-869 38.6 186.3 508.0 >1000
15 869-976 60.0 289.6 292.1 508.0
16 976-984 58.7 283.3 330.2 508.0
17 984-1139 68.5 330.6 254.0 355.6
TABLE 2 - FLAW SIZES FOR SPECIFIED WALL THICKNESS
AND SHUT-IN HOOP STRESS
Location Stress Flaw Length, mm
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Yield Strength Tensile Strength
Minimum: 483 MPa 565 MPa
Maximum: 586 MPa 793 MPa
Inside Wall Hoop Net Internal
Case Diameter Thickness Stress Pressure
Shallow-Water: 610 mm 38.8 mm 331 MPa 422 barg
Deep-Water: 610 mm 44.0 m 106 MPa 152 barg
TABLE 4 - CASE CONDITIONS
Pipe: Grade X70
Increment KP % SYMS MPa T.W. d/t = 0.5
1 0-29 118.3 571.3 25.4 38.1
2 29-42 114.8 554.2 25.4 38.13 42-56 99.8 481.9 88.9 101.6
4 56-68 81.7 394.4 139.7 190.5
5 68-278 66.8 322.6 203.2 330.2
6 278-282 49.1 237.2 292.1 647.7
7 282-535 47.3 228.6 304.8 736.6
8 535-611 44.9 216.6 330.2 762.0
9 611-617 58.0 280.0 241.3 457.2
10 617-742 55.8 269.6 254.0 469.9
11 742-755 65.0 313.8 215.9 355.6
12 755-788 63.8 308.2 228.6 368.3
13 788-854 66.5 320.8 203.2 342.9
14 854-869 76.7 370.4 165.1 215.9
15 869-976 65.7 317.2 203.2 330.2
16 976-984 107.6 519.2 50.8 63.5
17 984-1139 118.3 571.3 25.4 38.1
Location Stress Flaw Length, mm
TABLE 3 - FLAW SIZES FOR HYPOTHETICAL 21.9-MM WALL PIPE
SUBJECTED TO OIP DESIGN PRESSURE
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Weld
Yield Strength,
MPa
Tensile
Strength, MPa
Elongation in
50 mm, %
Tensile
Strength, MPa
RANGE 492-536 593-642 53-59 635-637
AVERAGE 515.8 616.4 56.9 658.8
Base Metal
TABLE 6 - TRANSVERSE TENSILE PROPERTIES
711-MM O.D. x 41-MM WALL TRIAL PIPE
Element,
Wt. % Min. Max. Avg. Min. Max. Avg.
Carbon 0.06 0.08 0.07 0.08 0.09 0.08
Silicon 0.30 0.36 0.34 0.23 0.25 0.24Manganese 1.58 1.64 1.62 1.61 1.66 1.64
Phosphorus 0.009 0.011 0.010 0.010 0.011 0.010
Sulfur 0.001 0.001 0.001 0.001 0.001 0.001
Aluminum 0.032 0.043 0.039 0.038 0.046 0.043
Copper 0.16 0.20 0.17 0.02 0.03 0.03
Chromium 0.03 0.03 0.03 0.02 0.04 0.03
Nickel 0.22 0.39 0.28 0.20 0.22 0.21
Molybdenum 0.00 0.02 0.01 0.01 0.01 0.01
Vanadium 0.07 0.08 0.08 0.08 0.08 0.08
Titanium 0.02 0.03 0.03 0.02 0.02 0.02
Niobium 0.038 0.043 0.041 0.043 0.051 0.046
Nitrogen 0.0030 0.0050 0.0039 0.0029 0.0038 0.0034
C.E. 0.37 0.40 0.39 0.39 0.41 0.40
Pcm 0.17 0.20 0.19 0.18 0.20 0.19
Plate Mill A Plate Mill B
TABLE 5 - CHEMICAL COMPOSITION
711-MM O.D. x 41.0-MM WALL TRIAL PIPE
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TABLE 7 - 711-MM O.D. x 41-MM WALL TRIAL PIPE
CVN Tests at -10 deg. C
(Average of 3 Specimens)
Base Metal Weld
Joules % Shear Joules % Shear
RANGE 216-321 100 143-181 96.7-100
AVERAGE 284.00 100 159.00 98.8
DWTT Tests at -10 deg. C
Energy, KJ % Shear
RANGE 18.3-42.0 90-100
AVERAGE 27.9 95.3
CTOD Tests at -10 deg C
CTOD, mm
RANGE 0.373-1.559
AVERAGE 0.95
MMAXIMUM HOOP STRESS VS. LOCATION
Fig. 1
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FLAW SIZES FOR SPECIFIED W.T.
CHARPY UPPER SHELF ENERGY = 100 J
0
100
200
300
400
500
600
700
800
900
1000
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17
Pipeline Route Segment
Flaw
Length,m
m
Through Wall Flaw 50% Surface Flaw
Fig. 2
FLAW SIZES FOR SPECIFIED W.T.
CHARPY UPPER SHELF ENERGY = 100 J
0
100
200
300
400
500
600
700
800
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17
Pipeline Route Segment
Flaw
Length,mm
50% Surface Flaw
Through Wall Flaw
Fig. 3
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0
0.2
0.4
0.6
0.8
1
FLAWD
EPTH/W
ALLTHICKNESS(d/t)
0 50 100 150 200 250 300 350
CRITICAL FLAW LENGTH, mm
INTERNAL FLAW, CTOD=0.38mm INTERNAL FLAW, CTOD=0.64mm
EXTERNAL FLAW, CTOD=0 .38mm EXTERNAL FLAW, CTOD=0 .64mm
SHALLOW-WATER CASE
WELD METAL
Fig. 4
0.2
0.4
0.6
0.8
1
FLAWD
EPTH/WALLTHICKNESS(d/t)
0 50 100 150 200 250 300 350 400
CRITICAL FLAW LENGTH, mm
CTOD=0.38mm CTOD=0.68mm
DEEP-WATER CASE, WELD METAL
INTERNAL FLAW
Fig. 5
0.2
0.4
0.6
0.8
1
FLAWD
EPTH/WALLTH
ICKNESS(d/t)
0 50 100 150 200 250 300 350 400
CRITICAL FLAW LENGTH, mm
CTOD=0.38mm CTOD=0.64mm
SHALLOW-WATER CASE, BASE METAL
INTERNAL FLAW
Fig. 6
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50
100
150
200
250
300
350
400
PRESSURE.BARG
0 100 200 300 400 500 600 700VELOCITY, M/SEC
688-MM O.D. x 38.8-MM WALL GRADE X70
SHALLOW-WATER CASE, CVN = 45 JOULES
Fig. 7
0
25
50
75
100
125150
175
200
DIFFERENTIALPRESSURE
.BARG
0 100 200 300 400 500 600 700 800 900 1000VELOCITY, M/SEC
698.5-MM 0.D. x 44.0-MM WALL GRADE X70
DEEP-WATER CASE, CVN = 3.4 JOULES
Fig. 8