propagación de fracturas tuberías de gas

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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


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


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

propagación de fracturas tuberías de gas

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