1-s2.0-0308016194001208-main

ELSEVIER

ht. J. Pres. Ves. & Piping 61 (1995) 433-456 030%0161(94)00120-0 Elsevier Science Limited

Printed in Northern Ireland 0308-0161/95/$09.50

FRACTURE MECHANICS ASSESSMENT OF INDUSTRIAL PRESSURE VESSEL FAILURES

N V Challenger, R Phaal and S J Garwood TWI, Abington Hall, Abington, Cambridge, UK

ABSTRACT

Fracture mechanics assessment procedures, such as BSI PD6493:1991, R6 and ASME XI, have become well established in industry. These published procedures provide methods for assessing the acceptability of flaws in fusion welded structures. For such procedures to be used with confidence, it is essential that their application be validated by comparison with large scale fracture mechanics tests, and actual structural failures. This paper describes eight industrial pressure vessel failures, for which PD6493 fracture assessments have been performed. It has been demonstrated that the assessment procedures are safe to use, provided that input data are reliable.

INTRODUCTION

Fracture is an important consideration when evaluating the integrity of welded structures. Pressure vessels and related systems form a class of components for which particularly high levels of integrity are required. This is due to the potential hazards which are associated with many industrial processes, combined with their high capital value. Pressure vessel failure can cause extensive damage to plant, owing to possible explosion and fire. Many pressure vessel failures are due to brittle fracture, where the tremendous potential energy stored in high pressure systems can lead to extensive fragmentation of the vessel, and high projectile velocities.

Fracture mechanics has proved to be a powerful tool for managing the risk associated with pressure vessel failure, as well as other welded components. Fracture assessment procedures such as BSI PD6493: 1980 (1) and R6 (2) have become well established in industry. They provide a rational basis for the safe assessment of welded joints, based on fracture mechanics principles. The fracture mechanics section of PD6493 has been extensively

433

434 N. V. Challenger et al.

modified by BSI committee WEE37, leading to a revised version being published in 1991 (3), and further modifications are under consideration.

For procedures such as PD6493: 1991 to be used with confidence for the assessment of fracture critical components, it is essential that they be validated by comparison with results generated from structurally relevant tests, such as wide plate, pipe bend and pressure vessel experiments, as well as with actual industrial failures.

Since the publication of PD6493: 199 1, an extensive validation programme has been undertaken, covering over 200 large scale fracture mechanics tests performed at TWI in the past (4), and published in the literature (5), as well as a range of industrial failures. This paper describes eight catastrophic pressure vessels failures (summarised in Table l), and presents the results of PD6493 fracture mechanics assessment of the failures.

PD6493 PROCEDURES

The fracture section of PD6493:1991 incorporates three levels of fracture assessment. Level 1 is similar to PD6493:1980, although assessment results are expressed in terms of a two- parameter failure assessment diagram (FAD), illustrated in Fig. 1. The Level 2 FAD is derived from R6 Rev. 2 procedures, while the Level 3 FAD is based on R6 Rev. 3. Elastic-plastic fracture is accounted for, as well as possible plastic collapse of the ligament. The interaction between these two modes of failure is accounted for by plotting a failure locus on the FAD. The abscissa, S,, represents a measure of plastic collapse of the ligament, while the ordinate, K, or dS,, ia a measure of fracture. The assessment of a specific flaw generates a point on the FAD. If this point lies within the failure locus then the structure may be considered safe. If the point lies outside the failure locus then structural failure is possible. If the point lies on the locus then the flaw may be considered to be critical.

1.2

Level 1 assessment line ,

1 UNSAFE

0.47 SAFE

0.21

0

0 0.2 0.4 0.6 0.8 I 1.2 1.4 1.6

Sr

Figure 1. PD6493 Level 1 FAD

Industrial pressure vessel failures 435

TABLE 1 Summary of failures reanalysed

Failure Initiating defect Causes of failure Circumstances of failure

Exxon pressure vessel - Port Jerome (1981)

Union Oil amine absorber (1984)

Vertical refinery tower (1981)

Typpi Oy ammonia plate cooler (1970)

Cockenzie power station boiler (1966)

John Thompson (1965)

Robert Jenkins (1970)

Ammonia catchpot (1982)

Cracks in PV shell at attachment weld root

Hydrogen cracking in HAZ of repair weld

Transverse weld metal hydrogen crack extended by creep

Fabrication defect extended by stress corrosion cracking

Arrested brittle crack at nozzle weld/drumshell, ori- gin unknown (present prior to PWHT)

Arrested brittle fracture from embedded transverse HAZ hydrogen crack

Arrested brittle crack from liquation crack extended by hydrogen cracking

Hydrogen cracks at fillet weld toe on vessel inner wall

Residual stress from attachment weld and strain ageing embrittlement

Hydrogen embrittlement and residual stresses

Low fracture toughness of plate and weld metal at hydrotest temperature, high residual stresses

Insufficient PWHT - high residual stress, low fracture toughness

Large initiating defect at a stress concentration

Insufficient PWHT - high residual stress low fracture toughness

Low fracture toughness, high stresses

Very low fracture toughness

Standby condition - pressurised to 75% op. pressure at ambient temperature

Normal operation

Hydrotest

Normal operation

Hydrotest

Hydrotest

Hydrotest

In service

Level 1 is the most simple and conservative of the three levels, incorporating a nominal safety factor of two on flaw size in terms of fracture, and 1.25 on load in terms of plastic collapse. Level 2 (see Fig.2) provides a more comprehensive and accurate assessment procedure. The primary and secondary stress distribution near the flaw under consideration is accounted for in more detail, as well as local stress concentration factors such as fillet weld


toes. A plasticity correction factor is included to account for crack tip plasticity in the presence of secondary stresses. There are no explicit safety factors built into Level 2, although guidance is provided in Appendix A of PD6493 for the application of partial safety factors. The conservatism of the assessment procedures is ensured by careful selection of appropriate lower bound values of material properties (fracture toughness and tensile properties), and upper bound stress distributions.

1.2 / I

, Level 2 assessment line ! ,,/

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8

Sr

Figure 2. PD6493 Level 2 FAD

Level 3 (see Fig.3) is the most comprehensive and complicated assessment procedure. Ductile tearing is accounted for, and knowledge of the material toughness is required in the form of an R-curve. The failure locus is material dependent, and the material stress-strain curve should be known. In cases where this information is not available a conservative default curve has been established (Eq.[25] of PD6493). This is particularly relevant to assessment of heat affected zones, for which the stress-strain curves cannot be accurately determined. The assessment of a specific flaw generates an assessment locus. If the locus lies entirely outside the FAD then tearing is predicted to occur, followed by unstable fracture. If the assessment line cuts the failure locus then the tearing is predicted to stop, and the flaw may be considered to be safe. The critical flaw condition is defined by an assessment locus which just intersects, or is tangential to the failure locus.

1.6

T 1.4- P s 1.21 -6 1 c2 l- 8 I

2 0.87

0.67

0.4-G

0.26

Assessment line - unsafe /

I ,!k-

Material specific FAD

Assessment line - safe

0 I /

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6

Lr

Figure 3. Typical PD6493 level 3 FADS


Normally, when performing fracture assessments to PD6493 procedures, the structure may be considered safe if the assessment point lies within the failure locus of the FAD, as described above. A point which lies outside the failure locus is of concern. In this report, since the fracture conditions of the pressure vessels are being predicted using PD6493 procedures, the assessment points are expected to lie outside the failure locus if the method is conservative. Points lying inside the failure locus are of concern, and represent non- conservative analysis procedures.

It should be noted that PD6493 is not a prescriptive document, and has not been standardised. Rather, PD6493 is an evolving document, providing guidance on the assessment of fusion welded joints, and there are many valid interpretations of PD6493.

In this study, each failure has been assessed using the procedures recommended by PD6493. The input data used for each assessment has been obtained from the published failure investigation reports. In accordance with PD6493: 199 1, apparently conservative assumptions have been used for all assessments; for example, lower bound toughness values and yield magnitude residual stress assumptions. However, because the assessments have been performed with the benefit of hindsight, a range of assessments has usually been performed for each failure. These use varying assumptions, some of which are subsequently shown not to be conservative and result in unsafe predictions from which various lessons can be learnt.

Due to the obviously unexpected nature of real failures, the quality of the input data is often lower than would be expected from a laboratory test. Where data has been unavailable, or of dubious nature, this has been indicated, and suitable engineering judgements made. This process is much closer to the typical situation where an assessment is being performed on an existing or postulated defect, with usually very restricted input data; therefore the assessments performed and described in this report are representative of many assessments which are performed for existing plant.

The PD6493: 199 1 assessments described in this paper have been performed with the aid of two computer programs, developed by TWI (Crackwise and PC6493). These programs have been validated (6) to ensure that the fracture assessment procedures of PD6493:1991 are implemented correctly.

ASSESSMENT OF PRESSURE VESSEL FAILURES

Port Jerome pressure vessel This failure (7) occurred in a 20 year old carbon steel spherical pressure vessel. The vessel was one of five identical reactors in a cyclic catalytic reforming unit, 4.3m in diameter, fabricated from 41mm thick plate (with no post-weld heat treatment). The failure occurred while the reactor was in a standby condition and so was pressurised (24barg, 75% of operating pressure) at approximately 5C with treat gas (primarily hydrogen).

The failure analysis concluded that failure initiated from cracks (primarily lamellar tearing) in the pressure vessel shell at the root of the weld used to attach a shroud support ring. It was concluded that the plate at the defect tips was embrittled by strain-ageing during


the welding of the shroud support ring. This support ring had been changed a number of times and the ring at the time of the incident had been welded to the pressure vessel shell approximately seven months previously.

The exact dimensions of the initiating flaw are not known. However, surface cracks in the vessel shell, extending from the attachment weld root, were found up to 3mm deep. Examination of the fracture surface indicated that the failure had initiated adjacent to a significant length of this attachment weld, extending to a through-thickness crack which deviated away from the attachment weld. Therefore the initiating defect is thought to be a surface defect of depth a up to 3mm and of an unknown but much greater length.

Residual stress was assumed to be uniform and of yield strength magnitude, relaxed in accordance with PD6493 procedures. Applied stress (69MPa, with stress concentration factor SCF=1.6) were based on finite element analyses reported in Ref.7, with M, factors recommended by PD6493 used for comparison, where M, is the stress intensity factor magnification factor, to account for the stress concentrating effect of weld toes.

Minimum values have been used for parent plate yield and tensile strengths, in accordance with PD6493 procedures. Three different levels of toughness have been assumed for all assessments. A value of crack-tip opening displacement (CTOD) toughness of 6,,=0.42mm represented the lowest value of toughness measured from full thickness CTOD specimens (at 5C) for as-received parent plate. Lower bound values of S,, = 0.03 and O.Olmm were reported in Refs. 7 and 8, respectively, for full thickness strain-aged CTOD specimens.

Since the actual flaw lengths were not reported in Ref. 7, maximum tolerable flaw depths have been calculated as a function of flaw length, for comparison with the reported 3mm flaw depth. Six Level 1 and 2 assessments have been performed, for three levels of toughness (Fig.4), based on the SCF data presented in Ref. 7. Three additional assessments have been performed using weld toe M, factors recommended by PD6493 for comparison (Fig.5).

IS

60 90

Crack half ten@. c. m m

Figure 4. Port Jerome pressure vessel - Level 1 and 2 assessments


30 60

h4kFrtwwd.SCF=l.O Mk fwor DM used. SCF - 1.6

Q.

5

- %TODmii = 0.b3fG 0 kk, CTODmat=OOlmm --i------.-r- ~_~~ ., i -- . ~. / I 0 30 60 90 120 I50

Crack half length, e, m m

Figure 5. Port Jerome pressure vessel - Level 2 assessments - effect of SCF

It can be seen that use of the as-received parent material toughness results in non- conservative assessments. This is not surprising, as the failure investigation identified that as- received parent plate toughness was not low enough to explain the brittle fracture at such a low load (75% of operating pressure).

Analyses performed using the fracture toughness values from the strain-aged CTOD specimens do conservatively predict failure, as can be seen from Figs 4 and 5. It should be noted, however, that the degree of strain-ageing in the actual pressure vessel is not known relative to the strain-aged CTOD specimens. The strain-aged CTOD specimens were designed to maximise the strain-ageing effect, by welding over prepared notches of a depth of 5mm to 6mm, which was found to give the greatest degree of embrittlement. However, it is considered that this technique is likely to realistically simulate the embrittlement seen at the initiating defect caused by the attachment weld.

It can be seen from Fig.4 that in this case Level 1 analyses are significantly less conservative than the Level 2 approach, particularly at the larger tolerable crack depths associated with the highest level of toughness. This is due to the use of the CTOD design curve, as recommended by PD6493.

Union Oil amine absorber pressure vessel In 1984 an amine absorber pressure vessel exploded, causing 17 fatalities and extensive damage (9). The vessel was an 18.8m high, 2.6m diameter, 25.4mm thick, carbon manganese steel cylindrical tower, used to strip hydrogen sulphide from the propane and butane process stream. The tower operated at just above ambient temperature (38C) and at a low internal pressure (14barg) and hence primary stresses due to pressurisation were calculated to be low (35MPa).

The failure investigation concluded that the vessel failed adjacent to the repair weld between courses 1 and 2. Course 2 had been replaced in 1974 due to the discovery of hydrogen blisters and delaminations. The repair welds had not been stress relieved.


It was concluded that the initiating defect had developed during operation (probably by hydrogen cracking) in the hard brittle microstructure adjacent to the horizontal repair weld. This heat affected zone (HAZ) microstructure was caused by the repair procedure and was not tempered by subsequent passes close to the vessel surface. These surface defects had extended through the vessel wall (probably due to a hydrogen cracking mechanism), until the deepest crack extended through more than 90% of the wall thickness, when the remaining ligament failed by stable tearing, creating a leak. The through-thickness crack then propagated by slow tearing, until approximately 800mm long, at which point fast fracture occurred.

Fracture toughness values of S,,- -0.17 and 0.064mm were measured from conventional HAZ notched and hydrogen charged specimens, respectively (both at 38C). Conventional HAZ notched specimens tested at -40C gave a lower bound S,,=O.l lmm; -40C was considered to be the lowest temperature that may have resulted due to venting of the propane immediately before final unstable fracture. For Level 3 assessments, a tearing CTOD R-curve was assumed: 6=0.144Aa+O.O32mm, where Aa is crack extension. This R-curve offset power law fit was estimated from a figure in Ref.9 and was used up to the extent of available data (Aa=4.Omm).

A total of 10 assessments have been performed, summarised in Fig.6 and 7. It can be seen that all assessments which take account of residual stresses conservatively predict the failure of the vessel, irrespective of whether fracture toughness is obtained from conventional or hydrogen-charged CTOD specimens. Indeed, the safety factor in these assessments is considerable. The fracture mechanics calculations presented in Ref.9 concluded that CTOD obtained from hydrogen charged specimens was required to adequately explain the unstable fracture. However this study took no account of residual stresses; levels of hydrogen content in the steel at failure and in the CTOD specimens were also not known.

0 0.2 0.4 0.6 011 I I.2 1.4

7 7

6- 6

5- 5

p4- 4

3- 3

Qm-304Nhd

2- 2

I I

Qm=ON/mm' ..----_. .._..

0 -10 / I r-I---- ---T--.

0 0.2 0.4 0.6 0.8 I 12 14

Lx

Figure 6. Union Oil amine absorber pressure vessel - Level 3 assessments


0.6 0.8

0 0.2 0.4 0.6 08 I

Sr

Figure 7(a). Union Oil amine absorber pressure vessel - Level 1 assessments

0

7

6

5

$4

I

3 :

02 0.4 0.6 0.8 I 1.2 -.__

0 02 0.4 0.6 0.8 I 1.2

Sr

Figure 7(b). Union Oil amine absorber pressure vessel - Level 2 assessments

Vertical refinery tower This was a 13 year old, 26m high, 3.7m diameter cylindrical tower which failed in 1981 during a re-validation hydrotest, following the addition of two new nozzles (10, 11). The tower was manufactured in 24mm thick A204 Grade C, C-1/2Mo steel, clad with 3mm thick 405 stainless steel, and had been stress-relieved. The tower normally operated at 450C but the water temperature during hydrotest was 8C. Failure of the vessel occurred at a pressure of 2 lbarg (operating pressure was 1Obarg).

The failure investigation concluded that failure initiated at a weld between the main vessel cylinder head and a 2.3m high, 1.8m diameter boot at the base of the vessel. At this position, a compensating doubler plate was attached, resulting in a weld thickness of approximately 63mm. It was concluded that the initiating defect, which was clearly visible on the fracture surface, was a fabrication-induced, transverse, weld metal hydrogen crack


(51mm long by 38mm deep), which had propagated by a creep mechanism, due to the applied and residual stress field.

It was noted in Ref. 11 that the residual stress was unusually high for a stress-relieved vessel; this, together with locally high hoop stresses from the hydrotest, low fracture toughness at the hydrotest temperature, combined with a pre-existing defect, was sufficient to cause brittle fracture.

Lower bound weld metal properties have been used throughout the analyses, 6,,t = 0.019mm; however the fracture toughness of the compensating plate and vessel shell were not significantly higher (0.035mm).

Values of membrane and bending primary stress (P, and PJ have been obtained from Ref. 11. These were originally obtained by finite element analysis of the failed region; hence an SCF of 1.0 has been used for all analyses. The value of membrane residual stress (Q,) obtained from Ref. 11 was measured by the block removal and layering technique. This resulted in internal measured hoop residual stresses of up to 119MPa. Hoop residual stresses measured at the outer surface varied up to 171MPa. A value of 120MPa (residual stress at inner surface) was assumed for all assessments. In the absence of any measurements of residual stress, for transverse flaws in PWHT joints PD6493 recommends the use of a value equal to 30% of the room temperature weld metal yield strength. This results in an assumed value for Q,=l83MPa (i.e. > measured Q,).

Various approaches were used for calculating the applied stress intensity factor, K, in Ref. 11. Upper and lower bound values of K, 3145 and 2564Nlmm3, were calculated for the plate surface and deepest point of the crack, respectively. In addition the Newman-Raju K solutions in PD6493 were used, for the deepest point of the crack (2775N/mm32).

Assessment results are presented in Fig.8 and 9, where it can be seen that all assessments conservatively predict the failure of the vessel, irrespective of the K solution used, or the location of the K calculation (plate surface/deepest point).

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 08 09 SI

Figure 8. Vertical refinery tower - Level 1 assessments


I / I 7----.- --

0 0.2 0.4 0.6 0.8 I Sr

Figure 9. Vertical refinery tower - Level 2 assessments

Typpi Oy ammonia plant failure This failure occurred at the Typpi Oy ammonia plant, Finland, in March 1970 (12). A set of four 1. lm diameter high pressure gaseous effluent water coolers failed with no prior warning.

The failure was traced to an existing defect in the forged head chamber of heat exchanger B. The defect was an internal, circumferential, surface defect, located in the head chamber wall, at the toe of the weld overlay on the tube plate. The head chamber wall at this point is understood to have been 85mm thick, and the tube plate 270mm thick.

It is thought that the initiating flaw was a fabrication defect (3mm deep, 12mm long). This defect then extended by a stress corrosion cracking mechanism caused by residual water from the hydrotest, during the period before final installation (one year). The stresses responsible for this crack extension are thought to be residual stresses remaining due to an inadequate post-weld heat treatment (PWHT). The defect size at failure was approximately 5mm deep by 70mm long.

The quenched and tempered forged material in which the defect was situated was also found to have very poor fracture toughness, attributed to slow quenching, leading to formation of upper bainite.

Residual stresses of up to 137MPa were reported in Ref.12, measured using X-ray diffraction. This value was used in some assessments, while an assumed level of residual stress was used for others, assuming that PWHT had been successful (i.e. 15% of yield strength, in accordance with PD6493).

An applied stress (P,) of 116SMPa (18% of yield strength) is mentioned in Ref. 12 as the stress caused by the process pressure, and was used in some assessments. Based on the pressure at failure (23MPa), a hoop stress of 136MPa has been calculated using standard thick shell solutions. For Level 1 calculations an SCF of 1.48 has been assumed, based on the PD6493 M, solutions for flaws at weld toes.

444 N. V, Challenger et al.

Fracture toughness data were obtained from two SENB specimens close to full section thickness (76mm), giving valid plane strain KrC results. The lower bound value was used for all assessments (1316N/mm32).

The results of all assessments summarised Fig.10 and 11. It can be seen from Fig.10 that all Level 1 assessments are conservative. All Level 2 assessments resulted in conservative assessments, except where it was assumed that PWHT was effective at stress relieving, which was not the case (using the lower level of applied stress assumed).

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.2 1.2

.

1

t

. 0

-F 0

0.8

I

I .s

t I I r; 0.6 3.6

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 sr

Figure 10. Typpi Oy ammonia plant failure - Level 1 assessments

0 0.2 0.4 0.6 0.11 1.2

.

t . . I I I

0 0

O.S-- 0.8

I & 0.6 & 0.6--

0.4 0.4--

0.2 : 0.2-- 0.2

0 I I +.+----+- 0 -

0 0.2 0.4 0.6 0.8 1 Sl

O.,,.- 0 0.2 0.4 0.6

Sl

Figure 11. Typpi Oy ammonia plant failure - Level 2 assessments

These assessments illustrate that small variations in assumed values can be crucial to the conservatism of an assessment in certain cases. In borderline cases, it is clearly advisable


to perform sensitivity analyses and/or apply partial safety factors where any doubt exists over input data.

Cockenzie power station boiler failure This failure occurred at the Cockenzie power station, Scotland, in 1966 (13). The first of the power stations 1.65m diameter boilers was approaching its full hydraulic test pressure for the fourth and final, intended, on-site hydrotest, when the vessel failed by brittle fracture. The failure was subsequently analysed by Burdekin and Dawes (14) using contemporary fracture mechanics assessment principles.

The initiating defect was identified as an internal, surface breaking defect (89mm deep by 330mm long), adjacent to an economiser nozzle and an internal welded attachment bracket. The economiser nozzle had previously been replaced during drum manufacture, but Ref.13 concluded that there was no evidence to suggest that the nozzle replacement was responsible for the formation of the arrested brittle fracture. Post failure examination of this initiating defect revealed that it was an arrested brittle crack at the drum shell/internal nozzle weld interface, and was not associated with the bracket weld.

The fracture face of the initiating defect was coated with oxides, indicating that it had been present during the stress relief heat treatment. The failure investigation reported in Ref.13 could find no defect from which the initial arrested brittle crack initiated. No conclusions were drawn as to the cause of the arrested brittle crack.

The report (13) stated that there was no evidence of crack extension during the previous three hydrotests, which had all reached full pressure (4098psig) safely. The report offers no explanation why the final test should fail at the lower pressure of 3915psig; the ambient temperature of the final test is reported to have been 7C but no mention is made of the temperature of the previous tests. Even if the previous tests had been performed at warmer temperatures, there should have been a warm proof-stressing effect, leading to a minimum failure load equal to the previous hydrotest loads.

Fracture toughness was not originally measured; however Burdekin and Dawes (14) presented a fracture mechanics analysis of various failures, including the Cockenzie pressure vessel, and used a value of CTOD of 0.43mm. This value of fracture toughness was reported to have been obtained from nominally similar material from another casualty, though no further details are given. In addition a second value for fracture toughness of 0.255mm was obtained by Babcock and Wilcox Research Station, from specimens extracted from the casualty material, as was reported in a comment in Ref.14 by W.M.Ham. This lower value of fracture toughness was also used in a number of assessments.

Parent material yield and tensile strength data were used in all assessments, in the absence of any relevant weld metal data. This should be a conservative assumption for the calculation of stress ratio (S,), assuming that the weld is overmatching.

Residual stresses have been assumed to be equal to 15% of (parent material) yield strength, following the stress relief heat treatment. Applied hoop stresses were calculated using both thin- and thick-wall solutions (P,=174MPa; P,=160MPa, P,=lSMPa, respectively). The results of the latter calculation give a stress distribution which can be linearised over the


defect depth in accordance with PD6493. Clearly, the initiating defect is at a fairly severe stress concentrator (i.e. adjacent to both a nozzle and an attachment bracket.)., and so various levels of SCF have been assumed. In the absence of more specific guidance, an upper bound SCF of 3.0 has been assumed.

Assessment results are summarised in Fig.12 and 13, where it can be seen that at the higher level of fracture toughness assumed (6,,,=0.43mm), the level of SCF assumed is critical to the conservatism of the assessment. The lower value of fracture toughness (6,,,=0.255mm) would appear to be the more relevant of the two values, having been taken from the casualty material. Use of this lower value of toughness results in conservative assessments at both Levels 1 and 2, unless no account is taken of the SCF due to nozzle and attachment bracket.

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 I.8 I.8

.A l.6-- 1.6

Id-- 1.4

t '8

0 01 0.2 0.3 04 0.1 0.6 0.7 08 0.9 Sr

Figure 12. Cockenzie power station boiler - Level 1 assessments

0 0.2 0.4 0.6 0.8 2.2 A

2-- A 2

l.8-- I.8

l.6-- 0 I.6 A

8 l.4-- I I.4

rfi 1.2

I ,I

0.8--

0 0.2 0.4 0.6 0.8 I Sr

Figure 13. Cockenzie power station boiler - Level 2 assessments


There can be seen to be little difference, in this example, between the results obtained using applied stresses calculated assuming thin and thick walled pressure vessel equations; the simplification of assuming a thin wall is conservative in this case

John Thompson pressure vessel failure This was a large, thick-walled pressure vessel, manufactured by John Thompson (Wolver- hampton) Ltd. for use in an ammonia plant (15). The vessel failed by brittle fracture during hydrotest in December 1965. Despite being hydrotested, the energy released was sufficient to project several pieces; one 2 ton piece was projected 152 feet.

The fracture was identified as initiating from two similar sites in the HAZ of the submerged arc circumferential weld joining the large forging and the adjacent strake. The initiating defects were small, transverse, embedded cracks, located in the forging HAZ. These cracks were probably formed by hydrogen cracking due to the heavy segregation of carbon and alloying elements in the forging, locally increasing susceptibility to this form of cracking. Initiation was then judged to have occurred into the weld metal. Both initiation sites (referred to as 11 and I2 in Ref. 15) were very similar but I1 was judged to be the primary initiation site.

Defect 11 (depth 8.3mm, length 9.5mm, ligament height 14.3mm) was judged to have initially have been a small hydrogen crack from which an arrested brittle fracture was initiated, either following welding under the influence of residual stress, or during the early stages of hydrotesting. This resulted in a larger defect from which final failure was initiated.

A stress analysis of the region in which fracture was initiated was reported in Ref. 15, and the results of this have been used to provide primary stresses. This stress analysis predicted a maximum stress of 249.7MPa at an internal pressure of 5100 psig; the actual internal pressure at failure was reported as 5OOOpsig, and so this value of stress has been linearly scaled accordingly, to give the value of applied stress assumed (P,=244.8MPa). This is marginally higher than the value of hoop stress calculated using thin wall shell solutions (213.8MPa).

The vessel was reported to have undergone PWHT, however Ref. 15 concludes that the stress relieving heat treatment was performed at too low a temperature, resulting in inadequate stress relief, in addition to the very high weld metal yield strength and very low weld metal Charpy toughness. Residual stress (Q,,,) levels of yield (parent and weld metal) magnitude (relaxed by primary stresses in accordance with PD6493) have been assumed. Additional assessments have been performed assuming residual stresses relaxed by PWHT (30% of appropriate yield strength, in accordance with PD6493 for transverse flaws).

Assessments have been performed with varying assumptions concerning tensile properties, due to the position of the flaw in the HAZ region between two regions of very different tensile properties. PD6493 recommends the use of the lower bound tensile properties (ie, the parent forging properties) for all calculations, with the exception of the residual stress assumption. For the residual stress assumption PD6493 states that the appropriate yield strength is that for the material in which the flaw lies (for transverse flaws). Use of the overmatching weld metal yield strength is clearly the conservative approach to take in this case.


Results of all Level 1 and 2 assessments are summarised in Fig.14 and 15. It can be seen that using the value of fracture toughness given in Ref.14 results in assessments close to the FAD, both at Level 1 and 2, indicating that care is required with the choice of input assumptions. Use of parent material (or weld metal) tensile properties, and residual stress equal to weld metal yield strength, results in conservative predictions, as long as no allowance is made for residual stress relaxation due to PWHT. This approach follows the recommendations of PD6493 but relies on the knowledge that PWHT was ineffective; clearly a similar assessment, if it were to be made without the benefit of hindsight, would not necessarily be safe.

0.5 06 0.7 0.8 O/9,

. 1.4

1.2

I

0.8

0.6

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9

Sr

Figure 14. John Thompson pressure vessel - Level 1 assessments

0 0.2 0.4 0.6 0.8 I

1.6 1.6

1.4-- . 1.4

.

l.2-- I2 0

0 I I

&f 0.8--

0.6--

0.2

0 4 I I , 4 -+--.-.. 0

0 0.2 0.4 0.6 0.8 I

Sr

Figure 15. John Thompson pressure vessel - Level 2 assessments


It should be noted that the value of fracture toughness (K,,=1840N/mm3*) obtained from Ref.14 cannot be taken as wholly reliable, despite being described as a K,, value, as fracture toughness testing was in its infancy at the time. However, use of various Charpy-K,, correlations suggest that this is likely to be a conservative value. All correlations attempted using the lower bound Charpy result at test temperature (12J at 7C) resulted in values of K,, at least as great as the value of I&,t used. The most conservative result was achieved using a correlation suggested by Girenko (16), which resulted in a correlated value of K,, - = 1880N/mm3 .

Robert Jenkins pressure vessel This 35m long vessel failed by brittle fracture during an hydrotest in November 1970 (17). The vessel was a relatively thin walled, fabricated from medium strength carbon-manganese steel (ASTM 515 Grade 70 (1967), unlike the previous two vessels, which were constructed from thick section, higher strength steel. The vessel was constructed in two sections. The section in which failure occurred was 2.4m in diameter, 28mm wall thickness, and was not stress-relieved. The other section was larger, thicker, and was stress-relieved.

The failure investigation concluded that the failure had initiated from an external, axially oriented, surface crack (114.3mm long and 12.2mm deep), extending from the toe of a fillet weld attaching a compensating ring at a manway. It was proposed that this defect was an arrested brittle crack which had been created in the early stages of testing, and had initiated from a liquation crack that had extended by hydrogen cracking, under the influence of high residual stress and constraint.

The vessel failed at a pressure of 420psig. Primary stress (P,) has been calculated for this pressure, assuming the pressure vessel to be thin walled, which results in a value of P, = 130MPa. Because of the depth of the flaw, M, values were calculated to have decayed to 1.0 and so no stress concentration effect was assumed, for both Levels 1 and 2.

Two different assumptions have been made for residual stresses. The simplest assumption is for uniform (through the plate thickness) tensile residual stress equal to the plate yield strength (relaxed in accordance with PD6493, Q,=295MPa). The second assumption, also suggested by PD6493, specifically for defects at fillet weld toes, calculates a residual stress distribution based on the heat input of the adjacent weld run. Unfortunately details of the weld procedure for this fillet weld have not been published, and so a heat input of lkJ/mm has been assumed; this assumption results in a value for the depth of penetration (y) of the residual stress field of approximately 20mm, which is significantly greater than the defect depth. This assumption resulted in residual stresses of Q,=92MPa and Q,=203MPa.

Results of all assessments, to both Levels 1 and 2, are summarised in Fig.16, where it can be seen that all assessments conservatively predict the vessel failure. Use of the residual stress field, specifically for weld toe defects, suggested in PD6493, results in a slightly less conservative Level 2 assessment. Level 1 assessments are unaffected by this assumption, as residual stress is automatically assumed to be uniform (Q,,, + Q,).


0 0.2 0.4 0.6 0.0 I.8

1.6-- 0

1.4--

0. I I I I

0 0.2 0.4 06 0.0 I

Sr

Figure 16. Robert Jenkins pressure vessel - Level 1 and 2 assessments

Ammonia catchpot pressure vessel failure This failure involved a thick walled (62mm) ammonia catchpot pressure vessel (7m long, lm internal diameter), which failed catastrophically, by brittle fracture, in January 1982, after approximately 16 years of service. The failure resulted in extensive fragmentation of the vessel, and considerable damage to the surroundings up to 300m from the vessel site.

The failure investigation (18) concluded that the pressure vessel shell was extremely brittle at the service temperature (26C), demonstrated by the extensive fragmentation of the vessel. The fracture initiation site was identified as one of several pre-existing axially aligned surface defects at a fillet weld toe, on the inner wall of the vessel. These pre-existing defects were identified as hydrogen cracks.

The exact cause of failure, after such a period of service, could not be attributed to a specific cause; possible causes included a change in service conditions, hydrogen embrittlement of the parent steel in service and/or extension of pre-existing defects, possibly by a hydrogen assisted process or fatigue. However, it was concluded that the parent material was sufficiently brittle that very little change (service conditions, toughness or defect size) had been required to cause catastrophic failure. Essentially the vessel had been in a condition very close to fracture from the time it had entered service.

The vessel had undergone a stress-relieving heat treatment after fabrication, however the failure investigation (18) cast doubt on whether the final pass of the fillet weld of concern had undergone PWHT.

The initiating defect was an axially oriented surface flaw, located at the toe of a fillet weld on the inner vessel wall, close to mid-length in the vessel. The fillet weld attached a bracket to the inside of the vessel; the attachment length (L) used for calculating the M, factors is not given in Ref.18, but was estimated from photographs to be approximately 30mm. Assessments were also performed with the assumed attachment length L=40mm and with M, factors set to 1.0 (ie assuming no fillet weld), to act as a sensitivity study.


The approximate initiating flaw depth, a, is known to be 4mm, however flaw lengths are not given in Ref.18. The failure investigation report does include a photograph of the initiation region, indicating that the two possible initiating defects are long, relative to their depth (2c=30-40mm). In view of this uncertainty over initial flaw dimensions, calculations have been performed to calculate tolerable flaw dimensions for a range of flaws.

Two levels of assumed primary stress were used for the assessments. The lower level of primary stress assumed (P,=285MPa), was the hoop stress taken directly from Ref.18. Hoop stress estimated using the vessel dimensions, and assuming the pressure vessel to thin walled, was slightly higher (P,=314MPa). This latter value agreed closely with the value of hoop stress calculated using thick walled pressure vessel solutions at the inner surface. Residual stresses were assumed to have been reduced by PWHT in accordance with PD6493, to 15% of parent yield strength, despite the doubts cast on the condition of the weld by the failure investigation.

PD6493 does not distinguish between internal and external axial surface flaws, however it is known that bulging has little effect on internal flaws; therefore most of the assessments performed in this case do not include a bulging factor. One assessment has been performed with a bulging correction factor for comparison purposes.

Defect calculations have been performed to estimate (just) tolerable defect sizes for a range of defect dimensions. The calculated tolerable defect dimensions are recorded graphically in Figs.17 to 19 for Levels 1 and 2, where it can be seen that the only analyses which indicate that flaws of depth greater than or equal to 4mm are tolerable (i.e. potentially non-conservative assessments), are those which ignore the stress concentrating effect of the fillet weld; even then the assessments indicate that flaws of 4mm depth (or more) are only tolerable for very short crack lengths (2c


6 i 5 10 IS 20 2J 31

b...-+...--- -.+ __t__ .._ +...--- -.+ I5 20 25

Crack half Ien& Crack half Ien& c, m m c, m m

A8a&mcnt length. L = 3Omm

Figure 18. Ammonia catchpot pressure vessel - tolerable surface defect sizes - Level 2 (P,=285MPa)

0 J IO IJ 20 21 30 6

__.___-.-. P

I5 Crack half length. c, m m

Figure 19. Ammonia catchpot pressure vessel - tolerable surface defect sizes - Level 2 (P,=314MPa)

GENERAL DISCUSSION

It is clear from the above assessments that all non-conservative analyses are the result of incorrect input assumptions or poor input data.

Good examples of this include the assessments on the Port Jerome vessel and Union Oil amine absorber tower, both of which were subject to embrittlement that was only revealed post failure. Use of standard CTOD specimens in each case would have resulted in non- conservative assessments. The embrittling mechanism had to be simulated in the CTOD test specimens in order to achieve conservative predictions. In both cases the degrees of embrittlement in the structure was unknown, and hence the tests employed an arbitrary degree of embrittlement. In the case of the Union Oil amine absorber tower failure this was achieved


by saturating with hydrogen while in the Port Jerome vessel, locally intensified strain ageing embrittlement was simulated by welding over notches in parent material CTOD specimens.

Several pressure vessels (i.e. Typpi Oy ammonia cooler, John Thompson vessel and the vertical refinery tower) were described as being stress relieved, but in reality contained significant residual stresses which were concluded to be primary causes of failure. Again this fact was only revealed post failure in each case. Residual stresses are not frequently known, and hence usually have to be assumed, despite being of primary importance in most failures. PD6493 procedures for residual stress assumptions have been shown to be conservative for the cases examined, except in the cases mentioned above, where PWHT had been incomplete.

It is interesting to note that low fracture toughness was a feature of almost all the failures. Examination of the figures show that all assessment points/loci are located in the fracture dominated region of the FAD. By comparison, assessment points for the large scale fracture mechanics tests assessed previously (4,5) are located around the knee region of the FAD, or beyond the plastic collapse limit. It is clear that the majority of research effort has not being aimed at the region of the FAD (brittle fracture) where many catastrophic industrial pressure vessel failures occur, but rather at the region of the FAD where normal design conditions are appropriate.

It would appear that most benefit is to be obtained by the accurate definition of input data. This requires accurate flaw sizing, location and monitoring, better understanding of likely embrittling mechanisms, in addition to accurate initial fracture toughness measurements, and better understanding of the applied and residual stresses. Most of the failures illustrated here have demonstrated inadequacies in knowledge in one or more of these areas. Further refinement and reduction in conservatism of fracture assessment procedures are not warranted without a high degree of confidence in the input data. The cost of generating improved input data must be balanced by an appreciation of the risk and consequences of pressure vessel failure.

Improved quality assurance (QA) procedures, at all stages of a pressure vessels life (design, fabrication, service and maintenance), can lead to improved input data, for purposes of vessel life extension and failure investigation. Fairly small expenditures of money (for instance, storing test samples of material from welding trials) can have a significant impact on the confidence (and hence conservatism) of future structural integrity assessments. The advantages of a successful structural integrity assessment can be very high, if plant life extension or change of service is anticipated, or if litigation occurs following a failure.

Smith (19) has reviewed the causes of over 66 pressure vessel failures, of which 26% occurred during hydrotest, 20% were described as brittle fractures, 14% were attributed to creep, 20% were caused by H,, stress corrosion cracking or related phenomena, while in 7% fatigue was reported to play a role in failure. In 6% of cases inadequate stress relief treatment was applied to vessels, while 5% of failures occurred at weld repairs; 12% of failures were attributed to low toughness or manufacturing defects. In many of the above failures, considerable damage to plant occurred, with extensive fragmentation (in one case pieces weighing over 100 tonnes were thrown more than 400m).


There are clearly many interacting mechanisms which contribute to pressure vessel failures. It is important that engineers concerned with pressurised systems be aware of the factors which affect the structural integrity of pressurised components, throughout the vessel life (design, fabrication and maintenance), including the influence of stresses, flaws, mechanical properties, and environment.

CONCLUSIONS AND RECOMMENDATIONS

Eight ferritic steel pressure vessel service failures involving fracture have been re-assessed following the BSI PD6493: 199 1 fracture assessment procedures. Assessments have been performed at Levels 1 and 2 in all cases, and at Level 3 where suitable input data exist. The fracture assessment procedures in PD6493 have been demonstrated to be safe for all the failures studied, provided appropriate input data are used.

The following points arise out of the case studies described in this paper.

(1)

(2)

(3)

(4)

Particular care should be given to ensuring that fracture toughness input data are relevant to the assessment. Possible embrittling mechanisms, such as hydrogen embrittlement and locally intensified strain ageing embrittlement, should be considered and accounted for in fracture toughness testing, if thought to be present.

The PD6493:1991 procedures for assuming residual stress levels are conservative, including appropriate relaxation methods, so long as post weld heat treatment (if performed) is carried out fully.

Further refinement and reductions in the conservatism of fracture assessment procedures are not warranted without a correspondingly high degree of confidence in the input data.

The use of more advanced methods (Levels) of analysis may not offer significant benefit where input data is uncertain.

ACKNOWLEDGEMENTS

The.authors would like to express their thanks to the UK Department of Trade and Industry, and the Industrial Members of TWI for funding the Core Research Programme. Thanks also to staff of the Engineering Department at TWI, especially M G Dawes, for providing advice and suitable references for this report.


(1)

(2)

(3)

(4)

(5)

(6)

(7)

(8)

(9)

(10)

(11)

(12)

(13)

(14)

REFERENCES

PD6493:1980, Guidance on some methods for the derivation of acceptance levels for defects in fusion welded joints. British Standards Institution, London, 1980.

Mime, I., Ainsworth, R.A., Dowling, A.R. and Stewart, A.T. Assessment of the integrity of structures containing defects - Revision 3. Central Electricity Generating Board, London, 1987.

PD6493: 1991. Guidance on methods for assessing the acceptability of flaws in fusion welded structures. British Standards Institution, London, 1991.

Challenger, N.V., Phaal, R. and Garwood, S.J. Appraisal of PD6493:1991 fracture assessment procedures, Part I: TWI data. TWI Report 7 158.1/93/762.02, March 1993.

Challenger, N.V., Phaal, R. and Garwood, S.J. Appraisal of PD6493:1991 fracture assessment procedures, Part II: Published and additional TWI data. TWI Report, to be published.

Booth, G.S., Garwood, S.J., Phaal, R., Hurworth, S.J. and Brown, P.L. Fitness-for- purpose assessment of weld flaws using micro-computer software. 4th Computer Technology in Welding Conference, Cambridge, 3-4 June 1992.

Merrick, R.D. and Ciuffreda, A.R. Brittle fracture of a pressure vessel - study results and recommendations. API 48th Mid-year refining meeting Session on unexpected material failures - refinery performance impaired, May 1983, Los Angeles.

Confidential TWI Report 27475/2.

McHenry, HI., Read, D.T. and Shives, T.R. Failure analysis of an amine-absorber pressure vessel. Materials Performance Vo1.26, No.8, pp. 18-24 August 1987.

Garwood, S.J. and Harrison, J.D. The use of yielding fracture mechanics in post failure analysis. Pressure Vessel and Piping Technology - A decade of progress - 1985 ed. C R Sundarajan ASME pp. 1043-1054.

Confidential TWI Report 2757 1,

Moisio, T. Brittle fracture in failed ammonia plant. Met. Const. and Brit. Welding Jnl, January 1972.

Report on the brittle fracture of a high pressure boiler drum at Cockenzie power station. South of Scotland Electricity Board, January 1967.

Burdekin, F.M. and Dawes, M.G. Practical use of linear elastic and yielding fracture mechanics with particular reference to pressure vessels. Proc. of Inst. Mech. Eng. conf., London, May 1971.


(15) Brittle fracture of a thick walled pressure vessel. BWRA Bulletin, Vo1.7, No.6, June 1966.

(16) Girenko, V.S. and Lyndin, V.P. Relationship between the impact strength and fracture mechanics criteria CTOD,, and K,, of structural steels and welded joints in them. Automatic Welding, September, 13- 19.

(17) Banks, B. Pressure vessel failure during hydrotest. Welding and Metal Fabrication, January 1973.

(18) Confidential TWI Report 2277913.

(19) Smith, T.A. A review of pressure vessel failure experience - some failure case studies. Vol Ml, SMIRT 8 (Structural Mechanics in Reactor Technology) Conference, Brussels, August 1985, 187- 193.

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