Fracture Control for the Oman India Pipeline T.V. Bruno, Metallurgial Consultants, Inc. Abstract This paper describes the evaluation of the resistance to fracture initiation and propagation for the high-strength, heavy-wall pipe required for the Oman India Pipeline (OIP). It discusses the unique aspects of this pipeline and their influence on fracture control, reviews conventional fracture control design methods, their 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 requirements. Introduction The Oman India Pipeline (OIP) will transport natural gas approximately 1100 km from Oman to India under the Arabian Sea, 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-in conditions) and the design temperatures are 0ºC minimum, 50ºC maximu m. The pipeline will be constructed with U-O-E pipe made from low-carbon, low-sulfur, microalloyed steel plate manufactured with thermo -mechanical process control (TMPC) including accelerated cooling. The specified mechanical properties are shown in Table 1. Because so much of the pipeline will be in deep water, the hoop stress of approximately 70 percent of the length of the pipeline will be less than 50 percent of the specified minimum yield strength (SMYS). Therefore, for most of the pipeline the potential for fracture will be much lower than for most pipelines. Figure 1 shows the maximum hoop stress vs. location along the pipeline. Principles of Fracture Control Design Fracture control design of pipelines requires that under the most adverse conditions: 1) the pipe has sufficient fracture toughness to tolerate small flaws without fracturing; 2) if the pipe ruptures from any cause, the fracture is ductile; 3) the steel has the capacity to absorb sufficient energy to arrest a ductile fracture, or crack arrestors are added. Considerable research on the behavior of pipelines sponsored 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 these methods by full-scale burst tests as well as their widespread successful application has shown them to be adequate within certain 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 limits. Nevertheless, as will be shown, the methods can be conservatively applied to evaluate resistance to fracture initiation and to give a reasonable estimate of resistance to fracture propagation. Fracture Initiation. AGA-Battelle Equations. The resistance to the initiation of ductile fractures can be evaluated for through-wall or partial-wall flaws using Equations (A-1) and (A-2) shown in the Appendix, which were developed by Battelle under AGA sponsorship. These equations give the size of a critical flaw, i.e., one that will cause a leak or rupture, as a function of the Charpy V-notch (CVN) toughness, the pipe size and grade, and the hoop stress. Similar equations have been developed for high-toughness, the pipe size and grade, and the hoop stress. Similar equations have been developed for high-toughness pipe for which fracture initiation is independent of the CVN toughness but Equations (A-1) and (A -2) were used because the results are conservative. As a first approach, critical flaw sizes for the OIP were calculated assuming a CVN fracture toughness of 100 J as specified for the longitudinal weld seam, as opposed to 200 J for the base metal, for conservatism. For convenience only through-wall (T.W.) and 50-percent wall surface flaws are considered. The pipeline has been divided into 17 increments by wall thickness for design purposes. As shown in Table 2 and Figure 2, the calculated critical flaw sizes are very large, ranging from 254 mm to more than 1000 mm. Equations (A-1) and (A-2) have been verified experimentally only for wall thickness up to 21.9 mm for the OIP and using the hoop stress based on the actual design pressure we can calculate critical flaw sizes within the wall thickness limits for which the equations have been verified experimentally. These values are very conservative because the assumed wall thickness gives a higher hoop stress than the actual hoop stress. Table 3 and Figure 3 show the calculated hoop stresses and critical flaw sizes based on a constant wall thickness of 21.9 mm. First consider the pipe from KP segments 3 through 15. The flaw lengths over this portion of the pipeline are orders of magnitude above the limits of detectability by ordinary inspection methods. Moreover, the assumed wall thicknesses are 39.2 percent (36.0 to 21.9 mm) to 50.2 percent (44.0 to 21.9 mm) less than the specified wall thicknesses and the hoop stresses are 1.4 to 2.2 times the actual maximum design stresses.

Next consider the pipe in KP segments 1, 2, 16, and 17. Even in these shallow-water areas the flaw sizes assuming a 21.9-mm wall thickness are relatively large and well within the limits of detectability. For these segments the wall thicknesses are 43.6 percent (38.8 to 21.9 mm) to 46.7 percent (41.1 to 21.9 mm) less than the specified wall thicknesses and the hoop stresses are 1.7 to 1.8 times the design stresses.

From the above it can be seen that even with conservative assumptions the OIP has adequate resistance to fracture initiation, based on the AGA -Battelle equations.

BSI PD 6493. Resistance to fracture initiation can also be evaluated using crack tip opening displacement (CTOD) and the method of British Standard Institute's PD 6493 : 1991, "Guidance on 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 flaw size was determined for the weld and base metal, and for internal and external surface flaws. The results were plotted as critical flaw length vs. depth (d) expressed as a fraction of the wall thickness (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 cusps in the curves are due to the formulas for calculating stress intensity; in reality the curves would be smooth.) As shown, internal flaws have a smaller critical flaw size than external flaws and are therefore more significant. For the lower CTOD value, the critical internal flaw length for deep flaws (>d/t = 0.40) is in the neighborhood of 20 mm and increases rapidly for shallower flaws. Figure 5 shows the results for the deep-water weld metal, internal flaw (the external flaw size, which is larger, is not shown). For deep flaws at the lower CTOD value, the critical flaw length is in excess of 30 mm.

The shallow-water base metal internal flaw case is shown in Figure 6. At the lower CTOD value, the minimum critical flaw length is about 40 mm. The deep-water base metal case gives even larger flaws and is not shown.

The critical flaw sizes for the weld metal are smaller than those for the base metal because PD 6493 assumes residual welding stresses for the former. Also, for the same design conditions, 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 can 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 the pipeline away from the opening in each direction. Under some conditions 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 conditions the fracture speed is fast enough for the crack tip to always lead the decompression wave of the pressure necessary to cause arrest and the 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 also 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 vs. pressure on the same curve. For a given pipe size and grade at a given operating pressure, the fracture velocity varies inversely with CVN upper shelf toughness. The fracture velocity curve has a "J” shape and levels off at a constant pressure that represents the fracture arrest pressure. The decompression curve is a function of the gas composition. When the CVN toughness is such that the two curves are tangent, fracture is unstable and will eventually arrest. When the curves are separated, the pressure quickly reduces to the arrest pressure and the fracture arrests quickly. When the curves intersect, the crack tip remains at a pressure sufficient to support fracture and propagation continues.

Curves were generated for a shallow-water and a deep-water case to determine the CVN toughness necessary to preclude long fractures. As shown in Figures 7 and 8, the required upper shelf energies for fracture arrest are:

Shallow-water Case: ~45 J Deep-water Case: ~3.4 J

The required toughness for fracture arrest is extremely low for the deep-water case and lower than might be expected for the shallow-water case. One reason for the low values for the deep-water case is the low hoop stress; because of water pressure the tensile hoop stress is only 22 percent of SMYS. Both cases are influenced by the fact that the gas composition and high pressure are such that the gas is very dense and tends to behave more like a liquid than a gas, and decompression waves travel faster than less dense gases. Crack Tip Opening Angel. The CVN test currently is the most widely used test to evaluate the resistance of pipelines to propagating ductile fractures. Recently a new approach utilizing the crack tip opening angle (CTOA) has been proposed. (5-7) With this approach, the fracture resistance of the pipe, termed (CTOA) c’ is compared to the driving force of the pressurized gas, termed (CTOA) max’ for a given pipeline design. The equilibrium 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 fracture tests using three-point bending specimens of two different ligament lengths and the value of (CTOA)max is determined using a computer program called PFRAC.(7)

Ten CTOA tests were run on samples of 660-mm O.D. x 41.3-mm wall test pipe with a yield strength of 478 MPa. This pipe had been produced from plate with similar chemical composition and processing as specified for the OIP. The (CTOA)max was determined based on the OIP design conditions. The average CTOA of the ten specimens was 11.7º compared to the calculated (CTOA)max of 3.3º. The fact that (CTOA)c was more than three times (CTOA)max indicates that fracture propagation is highly unlikely.

Two other CTOA studies were conducted. One used a parametric equation based on a number of PFRAC runs with a broad range of variables.(6) Inserting the OIP conditions in the parametric equations gave the following values for (CTOA)max:

Methane 3.8º Rich Gas 4.8º

These values are slightly higher but reasonably close to that given by PFRAC for the OIP conditions. The second study consisted of calculating an approximate CTOA from CVN and drop weight tear tests (DWTT) run on a number of test pipes produced for the OIP project. Data from ten test pipes gave a range of CTOA values from 9.4º to 32.4º, with an average of 19.4º. Although the use of CVN and DWTT tests is not as accurate as proper CTOA specimens, the results are consistent with the other tests that indicate a high degree of resistance to fracture propagation. Effects of Water Pressure. Water pressure has three effects that tend to mitigate against fracture propagation. The first is a reduction in hoop stress, which increases with depth. Secondly, water offers physical resistance to the outward movement of the flaps that form behind a propagating fracture, thereby slowing the fracture; this effect is similar to but greater than the effect of earth backfill on an onshore pipeline and does not increase with water depth. The third effect is due to an overpressure wave in the water surrounding the pipeline, which is caused by the exiting gas. The overpressure wave also slows the fracture speed. Burst tests 9 have shown beneficial effects of the overpressure at depths of 12 meters or more.

The effects of water pressure give added assurance that the potential for crack propagation is very low in the deep-water sections of the pipeline.

Evaluation of Ductile Fracture Behavior . To guard against long propagating fractures it essential that the pipeline operate above its ductile-to-brittle transition temperature and that the fracture absorb enough energy to arrest a fracture. These two properties are commonly evaluated with either the CVN test, the DWTT, or both. The CVN test is generally suitable for thin-wall pipe but the full- wall DWTT is preferred for heavy-wall pipe.

Ductile-to-Brittle Transition Temperature. Because the size and grade of the OIP pipe are outside the range of variables for which the correlation between DWTT or CVN tests and full-scale behavior has not been verified, West Jefferson tests were conducted by Europipe on two pipes made to the OIP specifications. Two tests were conducted in the fall of 1995, one of which confirmed that the DWTT accurately evaluated the transition temperature. In the second test the pipe leaked without rupturing; this test is to be repeated.

Energy Absorption. Typically, a plot of CVN or DWTT energy vs. temperature shows a maximum energy level at the lowest temperature at which the fracture appearance is 100 percent shear and no increase in energy at higher temperatures. This behavior gives a flat upper shelf to the transition curve. Some steels, particularly TMPC steels, show an increase in energy

beyond the lowest temperature for 100 percent shear, which gives a rising upper shelf. The rising upper shelf is accompanied by through-thickness separations on the fracture surface of the CVN or DWTT specimens. Some of the tests on OIP test pipe have exhibited separations and a rising shelf and some have not. To fully evaluate the fracture propagation resistance of pipe with a rising upper shelf generally requires full-scale burst tests of several pipe lengths welded together and pressurized with gas. This issue relates only to the length, but not the likelihood, of a rupture and there are other means of limiting the length of fracture should a failure occur, such as with crack arrestors. The cost of full-scale burst tests compared to the cost of crack arrestors justify such tests when considering a single project such as OIP .

Properties of Trial Pipe General. A trial production of one kilometer of 711-O.D. x 41-mm wall pipe was produced to the OIP specification to be used for various test purposes. The pipe was made from plate produced by two suppliers, each supplier utilizing five different heats of steel. Tables 5 and 6 summarize the chemical compositions and tensile properties respectively. As shown in Table 5, the chemical composition was similar for the two plate suppliers and the pipe showed good weldability as measured by the IIW carbon equivalent (CE) and the Pcm. The tensile properties were quite satisfactory. Vickers hardness (10-kg) surveys were made in the base metal and weld and heat-affected zone of the longitudinal welds. All the values met the requirement of 248 HV 10 maximum, and the large majority of readings was well below this value. Fracture Toughness. CVN, DWTT, and CTOD tests were made on pipes from each heat of steel. The test temperature was -10°C and the results are summarized in Table 1. All the base metal CVN specimens exhibited 100 percent shear with energy levels well in excess of the specified minimum. The weld-metal CVN specimens showed lower shear and energy values than the base metal, as expected, but all values were well above the minimum requirements. The heavy-wall pipe will be given a thermal aging treatment to increase the collapse resistance. As reported elsewhere9 this treatment has no significant affect on the CVN toughness.

The shear area of the DWTT specimens easily met the requirements and the absorbed energy, reported for information only, was quite high.

The initial requirements for weld CTOD values was 0.15 mm because pipe of this size and grade had not been previously produced and the minimum value that could be attained consistently was questionable. As a result of the trial run, the minimum value was increased to 0.40 mm; only one test, the minimum shown in Figure 7, was below 0.40 mm. Discussion The primary fracture control method is to design against fracture initiation. Work completed to date shows that the OIP pipe will have sufficient toughness to resist fracture initiation from small flaws, with critical flaw sizes much lower than the limits of detectability by non-destructive inspection. However,

because it is virtually impossible to assure resistance to fracture initiation from all causes, such as marine accidents and other low-probability occurrences, fracture propagation must also be considered. For the OIP, consideration of fracture propagation is secondary to consideration of fracture initiation and is an issue only in the shallow-water areas of the pipeline. Moreover, fracture propagation is principally an economic consideration relating to the 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 crack arrestors may be more practical. Conclusions . 1. Based on conventional fracture control technology using conservative assumptions, pipe produced to the OIP specification (will have adequate resistance to fracture initiation under the most adverse operating conditions. 2. Resistance to fracture propagation evaluated by conventional 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 fracture propagation by full-scale testing, the use of crack arrestors may be more cost effective. 3. Tests on a trial production of one kilometer of pipe showed fracture toughness well in excess of the minimum requirements 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., Pèro, 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 -Pipeline Research Committee, Sept. 26-29,1993.

7. Basically, G., Demofonti, G., Kanninen, M.F., and Venzi, S., "Step by Step Procedure for the Two Specimen CTOA Test," Pipeline Technology, II, 503.

8. Preston, R., "Improvement in UOE Pipe Collapse Resistance by Thermal Aging," paper OTC 8211 presented at the 1996 Offshore Technology Conference, Houston, Texas, May 6-9.

9. Bruno, T.V., "The Effect of Water Overburden on Ductile Fractures in Gas Pipelines," Doc. No. 9100-ALA -RD-L-1001, 1995.

Yield Strength, MPaTensile Strength, MPaHardness, HV 10

CVN at -10 deg. C

Energy, J: Base MetalWeld

% 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 68.5 330.6 254.0 355.62 29-42 62.9 303.6 292.1 431.83 42-56 60.5 292.0 292.1 495.34 56-68 40.2 194.0 457.2 >10005 68-278 31.6 152.5 558.8 >10006 278-282 21.9 105.7 736.6 >10007 282-535 22.4 108.1 736.6 >10008 535-611 27.8 134.2 673.1 >10009 611-617 27.7 133.7 533.4 >100010 617-742 33.1 159.8 558.8 >100011 742-755 32.4 156.4 584.2 >100012 755-788 41.8 201.7 431.8 >100013 788-854 46.2 223.0 381.0 812.814 854-869 38.6 186.3 508.0 >100015 869-976 60.0 289.6 292.1 508.016 976-984 58.7 283.3 330.2 508.017 984-1139 68.5 330.6 254.0 355.6

TABLE 2 - FLAW SIZES FOR SPECIFIED WALL THICKNESSAND SHUT-IN HOOP STRESS

Location Stress Flaw Length, mm

Yield Strength Tensile Strength

Minimum: 483 MPa 565 MPaMaximum: 586 MPa 793 MPa

Inside Wall Hoop Net InternalCase Diameter Thickness Stress Pressure

Shallow-Water: 610 mm 38.8 mm 331 MPa 422 bargDeep-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 38.12 29-42 114.8 554.2 25.4 38.13 42-56 99.8 481.9 88.9 101.64 56-68 81.7 394.4 139.7 190.55 68-278 66.8 322.6 203.2 330.26 278-282 49.1 237.2 292.1 647.77 282-535 47.3 228.6 304.8 736.68 535-611 44.9 216.6 330.2 762.09 611-617 58.0 280.0 241.3 457.2

10 617-742 55.8 269.6 254.0 469.911 742-755 65.0 313.8 215.9 355.612 755-788 63.8 308.2 228.6 368.313 788-854 66.5 320.8 203.2 342.914 854-869 76.7 370.4 165.1 215.915 869-976 65.7 317.2 203.2 330.216 976-984 107.6 519.2 50.8 63.517 984-1139 118.3 571.3 25.4 38.1

Location Stress Flaw Length, mm

TABLE 3 - FLAW SIZES FOR HYPOTHETICAL 21.9-MM WALL PIPESUBJECTED TO OIP DESIGN PRESSURE

Weld

Yield Strength, MPa

Tensile Strength, MPa

Elongation in 50 mm, %

Tensile Strength, MPa

RANGE 492-536 593-642 53-59 635-637AVERAGE 515.8 616.4 56.9 658.8

Base Metal

TABLE 6 - TRANSVERSE TENSILE PROPERTIES711-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 0.08Silicon 0.30 0.36 0.34 0.23 0.25 0.24Manganese 1.58 1.64 1.62 1.61 1.66 1.64Phosphorus 0.009 0.011 0.010 0.010 0.011 0.010Sulfur 0.001 0.001 0.001 0.001 0.001 0.001Aluminum 0.032 0.043 0.039 0.038 0.046 0.043Copper 0.16 0.20 0.17 0.02 0.03 0.03Chromium 0.03 0.03 0.03 0.02 0.04 0.03Nickel 0.22 0.39 0.28 0.20 0.22 0.21Molybdenum 0.00 0.02 0.01 0.01 0.01 0.01Vanadium 0.07 0.08 0.08 0.08 0.08 0.08Titanium 0.02 0.03 0.03 0.02 0.02 0.02Niobium 0.038 0.043 0.041 0.043 0.051 0.046Nitrogen 0.0030 0.0050 0.0039 0.0029 0.0038 0.0034

C.E. 0.37 0.40 0.39 0.39 0.41 0.40Pcm 0.17 0.20 0.19 0.18 0.20 0.19

Plate Mill A Plate Mill B

TABLE 5 - CHEMICAL COMPOSITION711-MM O.D. x 41.0-MM WALL TRIAL PIPE

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

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

Len

gth,

mm

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

Len

gth,

mm

50% Surface Flaw

Through Wall Flaw

Fig. 3

0

0.2

0.4

0.6

0.8

1

FLAW

DEP

TH/ W

ALL

THIC

KNES

S(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 CASEWELD METAL

Fig. 4

0.2

0.4

0.6

0.8

1

FLAW

DEP

TH/W

ALL

THIC

KNES

S(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 METALINTERNAL FLAW

Fig. 5

0.2

0.4

0.6

0.8

1

FLAW

DEP

TH/W

ALL

THIC

KNES

S(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 METALINTERNAL FLAW

Fig. 6

50

100

150

200

250

300

350

400 P

RE

SS

UR

E. B

AR

G

0 100 200 300 400 500 600 700 VELOCITY, M/SEC

688-MM O.D. x 38.8-MM WALL GRADE X70SHALLOW-WATER CASE, CVN = 45 JOULES

Fig. 7

0

25

50

75

100

125

150

175

200

DIF

FER

EN

TIA

L P

RE

SS

UR

E. B

AR

G

0 100 200 300 400 500 600 700 800 900 1000 VELOCITY, M/SEC

698.5-MM 0.D. x 44.0-MM WALL GRADE X70DEEP-WATER CASE, CVN = 3.4 JOULES

Fig. 8