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Jorge Hau, Antonio Seijas, Tim Munsterman, and Anelsy Mayorga

Capstone Engineering Services, Inc.

1505 Highway 6 South, Suite 250

Houston, Texas 77077

[email protected]

ABSTRACT

Several independent cases are described to illustrate the approach used for the analysis andevaluation of aging equipment for continued service, in the petroleum refining and chemical processindustry. In all cases described, the results were obtained by testing and metallurgically evaluating smallsamples extracted from the equipment. Charpy impact testing was performed to determine impact testproperties as a function of temperature, in one case to verify compliance with requirements and in theother case, as part of a temper embrittlement assessment. Fracture toughness Crack-Tip OpeningDisplacement (CTOD) testing was performed to obtain the fracture toughness required in theassessment. In the remaining cases the small samples were used to obtain tensile testing specimens andto perform direct metallographic evaluation to determine the nature and extent of the degradation thatthe steel suffered during service.

Keywords: Degradation Mechanisms, Material Properties, Fitness for Service, Assessment for BrittleFracture, Toughness, Graphitization, Wet H2S Cracking, Stress Assisted Corrosion.

INTRODUCTION

Pressure vessels in the oil refining and process industry usually age with time in service. Thedegradation mechanisms are multiple and depend on the type and quality of material of construction, theoperating conditions, the process fluid, operating history, and time in service. The current approach is toperform fitness for service assessment

1assuming worst case or lower bound values. Metallurgical

evaluation is performed in some other cases to determine condition and suitability for continued service.In both cases, it would be ideal to have material properties measured in the equipment or component.Traditional methods to extract samples for performing these mechanical testing are cumbersome and

often not practical. Usually the equipment needs to be taken out of service, weld repaired and even postweld heat treated (PWHT).

This paper discusses the evaluation performed using test results obtained from samples extractedfrom the equipment. This was done by using a device that cuts a small sample that leaves a smooth and

Evaluation of Aging Equipment for Continued Service

Paper No. CORROSION200505558

1

Copyright

©2005 by NACE International. Requests for permission to publish this manuscript in any form, in part or in whole must be in writing to NACE International,Publications Division, 1440 South Creek Drive, Houston, Texas 77084 77084-4906. The material presented and the views expressed in this paper are solelythose of the author(s) and not necessarily endorsed by the Association. Printed in U.S.A.

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shallow depression. Fitness for service local thin area calculations are done to establish if weld repair isneeded in the area that the samples are removed and if the areas can be left without repair. The sampleextraction from the equipment can be done on stream while the component is hot in operation and thedetails are given in the next section. Since the cut depth is only a fraction of the wall thickness, it couldbe used for most environments without much restriction.

SAMPLE REMOVAL TECHNIQUE

The technique uses a hemispherical shaped toothed saw type of cutter powered by an air motor. Itis a proprietary design fabricated in-house and the service is commercialized under the name “ScoopSampler”. The surface is lubricated and cooled using a water mist system with a high temperaturelubricant. This cooling system prevents sparks, allowing the device to cut the samples from operatingunits.

The device is held on to carbon steel and low alloy steel equipment using strong magnets. Usingthis device, samples can be cut to precision of ± 3.2 mm (0.125 in). The cut samples have semisphericalshape with height that can be adjusted within a range from 9 to 38 mm (0.375 to 1.5 in). The samplediameter depends on the depth of cut and can be between 50 to 95 mm (2 to 3.75 in). The samples canbe removed from vessels operating at temperatures up to 538°C (1000°F). This technique can also beused to extract samples from nozzles and piping, provided there is enough wall thickness.

This technique has been used to extract samples from storage tanks, reactors and other pressurevessels in refineries and chemical plants. The extracted samples have been used for a variety of testsfrom bulk metallographic examination, hardness profiles, tensile tests, bending tests, fracturemechanical tests, and Charpy V-notch impact tests. The test results have been used for mechanicalproperties in fitness for service and metallurgical assessments for continued service. This paperdiscusses several examples, employing this technique performed in different plants and locations. Allthe examples discussed below are independent of each other.

RESULTS

Brittle Fracture Assessment of Storage Tanks

We have assessed the brittle fracture susceptibility of two storage tanks made of carbon steel SA-283-C and built in 1968 using Charpy impact testing. The nominal wall thickness of these tanks is 26mm (1.037 in). Charpy V-notch impact test requirements are specified in construction codes such asASME

2as a function of the minimum specified yield strength of steel for pressure vessels and in API

Standard 6503

for welded steel tanks for oil storage. For the material of construction of the tanks underconsideration the minimum impact test requirement is 20 J (15 ft-lb). The critical exposure temperature(CET) for these tanks was –4°C (25°F).

The Charpy V-notch impact testing provides information as to whether the steel will exhibitadequate toughness. The absorbed energy not only decreases as the temperature decreases but alsocarbon steel experiences a transition from ductile to brittle fracture behavior at the transitiontemperature. Catastrophic failure of the carbon steel can occur when the temperature drops below thetransition temperature. Under brittle behavior of the carbon steel, the crack initiation and propagationoccurs more readily by cleavage fracture.

The minimum permissible design metal temperature (MDMT) for these two tank shells withoutimpact testing is 10°C (50°F), according to API 650. For tanks to operate at temperatures lower thanthis MDMT, it is required that the plates have to be tested to demonstrate adequate notch toughness.Thus, the Charpy V-notch impact test should be carried out to ensure that the carbon steel will haveadequate toughness at temperatures below the referenced minimum design metal temperature.

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Temperatures of –12°C and –1°C (10°F and 30°F) were chosen for the tests. Usually minimumdesign metal temperatures are 8°C (15°F) above the lowest one-day mean ambient temperature of thelocation where the tank is installed.

Figure 1 shows the six samples that were obtained, three from each tank. The sample and divotdimensions are described in Table 1. We performed calculations in accordance with API RP 579 usingdiameter and maximum divot depth of 63 mm (2.5 in) and 15.9 mm (0.625 in), respectively. Thesecalculations resulted in a minimum remaining thickness 8.9 mm (0.349 in) which was adequate

thickness for pressure retention. Thus, areas from which the samples were removed did not require anyrepair. However, they were removed from places intended for the installation of new nozzles and part of the objective of the test was to determine toughness before proceeding with the installation of thesenozzles. The chemical composition was determined for both tanks and found to comply withspecification. The microstructures obtained using optical microscopy on these samples was found to betypical of SA-283-C. A total of twelve half size Charpy specimens were machined from the samples thatwere extracted. Since these samples were small to carry out the Charpy impact tests, extension tabswere welded to the ends of the test material where the V-notch was machined.

The absorbed energy and the transition temperature depend on the size of the Charpy specimenand since half size specimens were used, the correlation given by API RP 579 was used

1. For half size

specimen the absorbed energy conversion to full size was achieved by multiplying by two. The

procedure is to determine the transition temperature and then add 11°C (20°F) to account for thetemperature shift in using half size specimens. The impact tests results at –12°C and –1°C (10°F and30°F) are shown in Figure 2.

The data from the Charpy tests revealed that the carbon steel samples taken from the lower courseof the two tanks exhibited a Charpy impact property that more than met the requirement of 20 J (15 ft-lb) minimum. Figure 2 indicates that the temperature corresponding to this requirement is much lowerthan even –1°C (+10°F). By extrapolation, the Charpy impact energy values indicated that thetemperature corresponding to 20 J (15 ft-lb) was at least below –15°C (5°F) for subsize specimens. Thiswould translate to an equivalent full size temperature of less than –4°C (25°F), which is the CET for thiscase. Therefore the criteria of the API RP 579 Level 1 assessment were satisfied and no furtherassessment was deemed necessary.

Temper Embrittlement Assessment of Reactors

A reactor in catalytic reformer service was to be removed from service in one refinery to be reusedin another location with severe winter condition. The reactor is 1976 vintage 5.25 in (133 mm) thick wall reactor made of 2.25Cr-1Mo steel internally weld overlay with stainless steel. The thickness at thehemispherical heads is 2

9 / 16 in (65.1 mm). The design pressure was 1700 psi (11.72 MPa) and the

design temperature was 850°F (454°C). It was necessary to assess the condition of the reactor forcontinued use in different weather. The greatest concern with this type of reactor material steels hasbeen the temper embrittlement involving the reduction of toughness as a result of long term exposure inthe temperature range of about 343° to 593°C (650° to 1100°F).

API RP 9344 specifies average values at –20°F (–29°C) for three Charpy V-notch test specimens,heat treated in accordance with this recommended practice not be less than 54 J (40 ft-lb) with no singlevalue below 47 J (35 ft-lb). However, the form U-1A manufacturer’s data report for this reactorspecified a Charpy impact value of 34/47 J (25/35 ft-lb) at a temperature of 10°C (50°F). This wasinterpreted as an average not less than 47 J (35 ft-lb) with no single value below 34 J (25 ft-lb). Theimpact test result requirement in the year this reactor was built (1976) is different from the current.Thus, it is necessary to assess the condition of the material for continued use.

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Prior to extracting the samples, the reactor was visually inspected internally and externally. All theinternal welds were inspected with dye penetrant testing. The welds outside were all inspected withcontrasting wet magnetic particle testing. Manual and automated ultrasonic inspection was used forinspecting girth and longitudinal welds. No crack indications were found. Field metallography was usedto verify the microstructure in seven different locations. The microstructures were found to be typical of 2.25 Cr-1Mo. Four samples were removed from the reactor, one from each shell ring (it has two rings),one from a longitudinal weld, and one from the top head (Figure 3).

According to Section 3 of API RP 579, if service condition may result in degradation of thefracture toughness with time of service, the Level 3 assessment should be applied. Guidelines have beenissued

5regarding the minimum pressurization temperature of reactors, depending on the year the steel

was manufactured. The reactor under consideration would belong to the second generation (1973-1980),for which a minimum pressurization temperature of 149°C (300°F) has been recommended. Byperforming Charpy impact testing it was possible to reduce some conservatism in this approach.

By welding extension tabs to the samples extracted it was possible to obtain a total of 22 half-sizeCharpy V-Notch specimens, six for each sample, except for the weld that gave only four. The results aresummarized in Figure 4, after conversion to full size values. These results indicated that the top headappeared much tougher than rings 1 and 2, and also the weld. The results obtained for both rings werefairly similar. The curves were fitted manually and they superimposed each other. The curve for the

longitudinal weld appeared slightly tougher but this appearance would depend on the scale chosen forthe ordinate. The difference can be hidden or exaggerated depending on the range chosen for theabsorbed energy in the ordinate. Statistic methods can be used to resolve this and determine if the curvefor the weld really represented a result higher than those obtained for the rings. To do this, the resultsfrom these three locations were statistically analyzed (Analysis of Variance) and it was found that thedifference was not significant as compared with the data scatter. The results varied even for the samematerial tested at the same condition and if this data scatter is considered, it would not be possible todistinguish if any particular experimental point belongs to either the rings or the weld in this set of data,unlike the result from the top head that was significantly different and was kept separately.

When it is statistically found that there is not any significant difference between two or moresamples, it is customary to group them into one single set of data. This is convenient because the testing

produces a more accurate result when it is based on a larger amount of data points. In this case the pointsthat corresponded to the two rings and the welds were found not to differ significantly from each otherand the data from these three sources were fitted with a common and single straight line. The associated95% confidence curves were derived, as illustrated in Figure 5. The fitted straight line with itsconfidence limits represents the reactor that consists of rings, heads, and welds. The result from the tophead was not included because it exhibited superior Charpy impact properties. The principle is that if thereactor were to fail because of lack of sufficient toughness, it would fail in the areas that exhibited lowerimpact properties, in this case either the rings or the weld.

The Charpy impact values obtained still meet the limits of 34/47 J (25/35 ft-lb) at a temperature of 10°C (50°F) specified originally by the manufacturer of this reactor. At this temperature the averageobtained was higher than 47 J (35 ft-lb) and there was not any single value lower than 34 J (25 ft-lb).

This was the requirement originally specified for the construction of the reactor and because the data didnot indicate any reduction in toughness of the material in comparison with these referenced values, it isnot possible to state that the steel has suffered any embrittlement since it was originally constructed.

The value of 23 ft-lb (31 J) was taken for defining the minimum allowable temperature, based onthe requirement for Charpy V-Notch impact test found in Figure UG-84.1 in ASME Code Section VIII,Division 1. If the fitted straight line was assumed in Figure 5, this would correspond to about –11°C(13°F). Adding 11°C (20°F) to account for the temperature shift correction, the temperature would be0°C (32°F). However, because of the scatter, the point corresponding to the lower 95% confidence curve

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was chosen instead. This resulted in –5°C (23°F) which by adding 11°C (20°F) translates to 6°C (43°F).This value became the permissible lower metal temperature limit or Minimum Allowable Temperature(MAT). The reactor could be critically exposed to lower temperatures than this, particularly duringwinter time. Therefore, the reactor did not satisfy the criteria of the API RP 579 Level 1 assessment forbrittle fracture.

Once the minimum allowable temperature was established, Method A of the Level 2 assessmentwas used to derive a safe envelope of allowable temperatures as a function of pressure. There would be a

risk of brittle fracture if the reactor was stressed above this safe envelope, Figure 6. The Method A of the Level 2 assessment is described in detail in API RP 579 and consists in calculating the primarymembrane stress using the design conditions and also various possible operating pressures lower thanthe maximum allowable working pressure.

A ratio was then obtained by dividing the primary membrane stress derived from different pressurelevels by the primary membrane stress corresponding to the design conditions. Since the pressure levelsused are lower than the maximum allowable working pressure, the ratio will always be lower than unity.The actual applied stress is reduced accordingly with pressure reductions. Section 3 of API RP 579provides a curve for the reduction in the MAT based on the excess thickness. With the obtained ratio inthe ordinate the temperature reduction given on the abscissa is obtained. This amount was subtractedfrom the original MAT of 6°C (43°F) and a new minimum allowable temperature was thus obtained for

that particular pressure. As the pressure level was reduced in the calculation, lower and lowertemperature became acceptable because the stress level is reduced.

The calculation was repeated for even lower pressure levels until reaching a condition where thecritical exposure temperature (CET) was equal to or less than the corresponding MAT for this particularpressure. The CET was defined as the lowest metal temperature at which the reactor could be exposedto, in this particular case, a coldest winter day of -21°C (–5°F).

The reactor really operates hot at temperatures which are always higher than 6°C (43°F). For aslong as the actual metal temperature is higher than the assumed MAT, the reactor can be pressurized upto its maximum allowable working pressure. The reactor could only experience internal pressure at ornear ambient temperature during start-up, shut-down, abnormal upset conditions or during tightness or

hydrostatic test. Therefore, under no circumstances the reactor was to exceed the pressure level given bythe envelope in Figure 6, at any given temperature. Any combination of pressure and temperature wasconsidered to be acceptable for as long as it could be located below the safe envelope.

The above was valid only if the reactor was kept flawless. In the presence of a crack, the reactormay still fracture in a brittle manner. The reactor was thoroughly inspected for flaws and did not showany cracks but provision has to be made for future service. Therefore, the analysis was taken one stepfurther by performing a Level 3 assessment, involving stress, flaw size, and material toughness, basedon Section 9 about Assessment of Crack-Like Flaws, in the API RP 579 document. Material toughnesswas derived using the Charpy transition temperature of 6°C (43°F), obtained from Figure 5 as referencetemperature, and Figure 6 was used to define temperatures of the analysis to estimate lower boundfracture toughness as a function of temperature in accordance with expression given in Appendix F of

API 579.

Several damage tolerance curves were plotted for various conditions. Figure 7 shows the damagetolerance curve assuming the worst scenario of having the full design pressure at 6°C (43°F) and a crack parallel to the longitudinal weld. Figure 8 shows the corresponding fracture tolerance signature curve forcrack-like flaw aspect ratio close to 6 (crack length divided by depth). Damage tolerance and fracturetolerance signature curves were derived for all other possible crack configurations. The fracturetoughness decreases with temperature but the maximum allowable pressure has to be reduced inaccordance with Figure 6. This is why the critical crack size increases with decreasing temperature.

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Knowledge of the critical flaw size allowed to determine appropriate inspection method to be usedto ensure there will not be any crack approaching critical size during future service. By complying withthe safe envelope of temperatures and pressures given in Figure 6 and implementing an inspectionprogram to identify cracks before brittle fracture.

Low-Cycle Fatigue Assessment of Coke Drums

We have evaluated two coke drums that had a history of bulging and cracking in the vicinity of

circumferential welds for continued use until the replacement drums are built and installed. These bulgesand cracks are due to the severe operational thermal cycling experienced over the life of these cokedrums. Thermal cycles occur during filling with hot product at ~900°F (~480°C) and during waterquenching immediately after the coke forms. These coke drums needed to be evaluated to avoid risk of failure. It is necessary to develop the level and frequency of inspection required for earlier detection of critical cracks. Cracking was determined to be the primary life limiting factor, the risk being the firesassociated with leaks developed in the drums. Low cycle fatigue (LCF) due to cyclic thermal stresseshas been identified as the main damage mechanism in coke drums. The drums were subjected to severalprevious evaluations, including finite element analyses to estimate stress level arising during thermalcycles.

Samples were removed from the peak and tail of a bulged area on a SA-204 Grade C carbon-

0.5Mo steel coke drum shell section. An additional sample was removed from a location without anyapparent bulging. Test samples were prepared to conduct CTOD tests, Figure 9. Material fracturetoughness and crack propagation data were obtained from these tests and the results were used toconduct a fatigue crack growth analysis.

CTOD test method6

was used to measure fracture toughness using notched specimens sharpenedby fatigue pre-cracking. Three-point bend specimens were used after precision cutting and welding thenecessary extension tabs. The tests were conducted at room temperature. This was a conservativeapproach because the coke drums operate at higher temperatures. Clip gage displacement controlledloading were used in the specimens tested. Two CTOD specimens were prepared from each scoopsample, except for the location that was not bulged, where only one specimen was obtained. The resultsare shown in Table 2.

A conservative fatigue crack growth analysis for constant amplitude loading was used to estimatethe number of cycles required to have a through-wall crack. The operating cycle of the drums wasassumed to take 32 hours (including cleaning). According to the experience the shell of coke drumsundergoes one significant strain cycle every operating cycle. The maximum strain is generallyassociated with the quenching period. For crack growth, only the positive part of the stress cycle isrelevant.

Crack growth data for the materials was obtained from the literature7,8,9

and these values werecompared with those obtained during the CTOD specimens pre-cracking. Fatigue crack growth curvesfor C-Mo and Cr-Mo steels are used to derive the parameters “C” and “m” of the Paris equation

10: The

values assumed were C = 2.35 x 10-9

and m = 3.64.

da/dN = C ∆Km

where:

da/dN is the crack growth rate (“a” crack depth, “N” number of cycles),

∆K is the cyclic stress intensity factor.

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Uniform primary stress was used and made equal to the maximum axial stress induced by aninternal pressure of 0.69 MPa (100 psig) in the top shell section of the drums. Two different cyclicstresses were assumed for the fatigue analysis. The first one, equal to 225 MPa (32.6 ksi), was themaximum axial stress reported in previous finite element modeling

11. In the analysis the actual

dimensions of the bulged areas were considered in the construction of the geometrical model. In thesecond case, the cyclic stress was assumed equal to 296 MPa (43 ksi), the minimum specified tensilestrength for SA-204 Grade C steel

2.

Figure 10 shows the fatigue crack propagation curves (depth and length of the crack versusnumber of cycles) for an initial surface crack 6.3 mm (0.250 in) deep, with and aspect ratio (depth

/length) equal six. Two curves are shown: the lower one corresponded to a cyclic stress of 225 MPa(32.6 ksi) and the upper curve, to a cyclic stress equal to 296 MPa (43 ksi).

The remaining life associated to crack propagation was set as the number of cycles required for theinitial defect size “a” = 6.3 mm (0.25 in) to grow through full thickness. This initial crack would gothrough the wall after 220 cycles for a cyclic stress of 225 MPa and 80 cycles for a cyclic stress of 296MPa, corresponding to the two cyclic stresses used. However, in the crack propagation curves a ratherrapid increase in growth rate was observed when the crack depth approached and exceeded half the wallthickness (Figure 10). To be more conservative the number of cycles to failure for each case was set tobe when reaching half wall thickness. In the worst case assumed (when cyclic stress was 296 MPa or 43

ksi), the initial crack (0.25 in deep) would reach half the wall thickness (7.9 mm = 0.31 in) after only 40cycles. The inspection interval must therefore be set at every 40 cycles. For a full 32 hours cycle, thedrums must be inspected every two months to avoid having leaking through wall cracks and the firepotential. The recommendation was to perform inspection every 32 cycles which corresponds twomonths to detect cracks before these are able to reach critical size, until the opportunity becomesavailable to replace these coke drums.

Graphitization in Carbon Steel

We evaluated carbon steel reactor heads, the vapor outlet piping and header on top of the reactor,and overhead line from fluid catalytic cracking (FCC) service operating at a pressure of 2.8 kg/cm

2(40

psig) and at temperature ranging from 530° to 535°C (985° to 995°F). In this unit the outlet tube from

two cyclones go through the top head to a vapor header that acts like an external plenum chamber of thecyclone system. Two vapor outlet short pipe sections join the vapor header to the reactor top head in aconfiguration similar to outlet nozzles but without any flange connection. The reactor overhead linegoing to the fractionator was included in the evaluation. The bottom reactor head was also examined,together with a carbon steel pup piece that represents the inlet nozzle to the reactor. This is welded to atransition piece made of 1.25Cr-0.5Mo steel and then to the cold walled carbon steel reactor riser. Theevaluation was based extensively on in-situ metallography and replication and on samples directlyremoved from the lower elbow of the reactor overhead line, making the change in direction from verticalto horizontal.

The reactor heads were fabricated in carbon steel SA-515 Grade 55 and the vapor outlet piping andheader were fabricated with carbon steel SA-285-C and installed in 1989. The reactor shell wasfabricated in 1.25Cr-0.5Mo steel, internally clad with stainless steel type 410. Graphitization wasdetected in the past in carbon steel components and the objective of the study was to further assess theextent of damage and its fitness for service.

Graphitization involves a change in the microstructure after long term operation in the 427° to593°C (800° to 1100°F) range. It may cause a loss in strength, ductility, and creep resistance. At thistemperature range, the carbide phase (cementite, Fe3C) in carbon steel is unstable and may decomposeinto graphite nodules. They can form at the low temperature edge of weld heat affected zone (HAZ),resulting in a row or band of aligned graphite nodules that may extend across the wall thickness and

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parallel to the weld seam. This form of graphitization can result in a significant reduction in load bearingcapacity and thus increase the potential for brittle fracture along this plane.

Some steels are much more susceptible to graphitization than others or remain unaffected even if they conform to the same steel specification. Exactly what causes some steels to graphitize while othersare resistant is not yet fully understood

12.

Metallographic examination performed in several places in the reactor heads suggested that

graphitization did not occur in the reactor heads. The wall metal temperature of the reactor heads isbelow the range of graphitization. The reactor heads are internally protected with 75 mm (3 in) thick refractory wall. The carbon steel overhead vapor header and piping do not have any internal insulatingrefractory so they were all affected. The reactor shell does not have internal insulating refractory and ismade of 1.25Cr-0.5Mo steel that is not susceptible to graphitization

12.

The vapor outlet tubes, header and reactor overhead line to the fractionator had graphitization. Thecase for the reactor overhead line to the fractionator was among the most severe. Four small sampleswere extracted from the elbow joining the vertical portion of this reactor overhead line to a horizontalsection. These samples were examined metallographically on planes in the direction of the wallthickness. This volumetric examination confirmed findings made with field metallography replicationand aided the interpretation of the microstructure and features seen on the replicas. The specimen for

bending test was extracted from the lower reactor overhead line elbow.

It is hard to prepare a sample with V-notch coinciding with graphite nodules to perform Charpyimpact test. Thus, we chose the bending test that has been used for evaluation of graphitization in pipingsystems

13. It is a procedure that consists in applying bending until a crack forms and becomes visible. At

this point the test is stopped and the bending angle is measured. A correlation was established to rate thedegree of graphitization: less than 15° bending was considered very severe; from 15° to 30° severe; from30° to 50° strong; from 50° to 90° moderate, and higher than 90° was considered to be mild.

The bending action was stopped after a crack was visible, Figure 11. The bending angle wasmeasured, following the procedure described above. The bending angle was about 60°, which placed itinto the category of moderate graphitization. Figure 12 revealed that a crack was formed by alignment of

graphite nodules, and the actual fracture occurred by tearing apart the sound material that existedbetween adjacent and aligned graphite nodules. The locations of the weld metal and the heat affectedzone (HAZ) are shown in Figures 11 and 12.

The graphite nodules distributed more or less randomly in the base metal are usually consideredharmless. The effect of graphitization is severe when the nodules align parallel to the HAZ. The steelmicrostructure was free from pearlite and carbide spheroidization, Figure 13. Graphitization and carbidespheroidization are competing mechanisms of pearlite decomposition. The dissolution of cementite(Fe3C) plates and the diffusion of carbon are common to both mechanisms. The rate of decomposition istemperature and time dependent. The difference is on whether the diffusing carbon will nucleate andprecipitate as spherical Fe3C particles or as graphite nodules. Graphitization may be predominant attemperatures below about 550°C (1025°F)14, although the process is highly sensitive to the steelchemistry, affecting not only the stability of the cementite (which acts upon the dissolution ease andrate) but also on the nucleation process.

The graphitization observed involved graphite nodules formation, without any carbidespheroidization. The density of nodules observed did not differ significantly from the amounts reportedin previous in-situ metallographic evaluations carried out in 1997 and 1999. Tensile specimens weremachined out of these samples to derive tensile properties. The result is shown in Table 3. The steelcomplied with SA-285 Grade C that specified

2tensile strength of 380-515 MPa (55-75 ksi) and

minimum yield strength of 205 MPa (30 ksi).

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The most common cases of graphitization observed in this evaluation was evidenced by thepresence of graphite nodules randomly and widely distributed in the base metal. This type of graphitization is not considered critical. Aligned graphite nodules extending along a band or line parallelto the HAZ are referred to as eyebrow graphitization and this represents the most critical cases. Thecases that were detected were subjected to repair by removing the affected material. Any new carbonsteel placed in this system would likewise be subjected to graphitization but after certain amount of timein service. As for the existing steel, further graphitization seemed unlikely because the process hasalready converted most of the available carbides.

The disappearance of perlite and carbide particles in carbon steel during prolonged service at hightemperature involves softening and loss of creep strength. However, there has not been any indication orevidence of creep damage in the parts and components that are exhibiting graphitization anddecarburization. Therefore, loss of creep strength was not considered to be the predominant degradationmechanism. In regard to tensile properties, the strength requirement of a Level 1 assessment wassatisfied and no further assessment was deemed necessary. The steel still complies with tensile strengthrequirement.

Because eyebrow graphitization is critical it was recommended to continue inspection until all theremaining strength welds were examined to detect these severe graphitization cases. Since weldability isnot a problem, weld repair can be done with relatively ease once these cases are detected and repaired.

The affected area could be ground or gouged. A whole ring or window could be removed and replacedby a new ring or insert plate. The repairs would not need any PWHT. Alternatively, if replacement is tobe done, upgrading to 1.25Cr-0.5Mo steel is recommended.

Wet H2S Cracking

We evaluated two small samples extracted from the inside surface of a storage sphere to examinecross sections through weld seams and determine the nature of the cracking observed. One sample wasabout 76 mm (3 in) in diameter and 19 mm (0.75 in) thick. The second sample was about 67 mm (2.625in) in diameter and 13 mm (0.5 in) thick. The specified material was carbon steel SA-516-70. This was arefrigerated sphere originally used to store propane and currently used for propylene. It was constructedin 1968 and was not PWHT originally.

Cross sections through both samples were prepared and examined. Figure 14 shows the crosssection through the cracks in one sample. Figure 15 shows cross section through the cracks found indifferent metallographic specimens removed from the other sample. A Vickers microhardness mappingwas done. This cracking was always associated with hard spots and cracks arrested when reaching softermetal. The cracking occurred in the HAZ region immediately adjacent to the last weld beads depositedon both sides of the weld. The cracking was transgranular and initiated in the area adjacent to the toe of the weld, where the hardness was the highest. The microstructure in this cracked area appeared eithermartensitic, bainitic or both. Presence of sulfur was detected within the fracture surfaces whenexamining in the scanning electron microscope (SEM), using energy dispersive analysis (EDS). Thesefeatures observed in the cracking suggested wet H2S cracking. Hard steel regions in non PWHT weldsare likely to crack if wet H2S corrosion occurs.

The problem with liquid petroleum gas (LPG) storage sphere cracking has been described byothers

15,16. When these spheres started to be subjected to wet fluorescent magnetic particle testing

(WFMT), many indications were reported, some attributed to fabrication and other to wet H2S cracking.Susceptible hard HAZ may crack due to hydrogen picked up during welding and fabrication or from wetH2S corrosion occurring from accidental excursions. The cracking and its features will look alike sinceno evidence will be left to distinguish whether the hydrogen is coming from hydrogen picked up duringwelding or H2S corrosion. Iron sulfide scaling and film found in the cracks provided the evidence for

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suspecting wet H2S cracking. The sphere has been used to store clean propane and propylene, whichunder otherwise normal condition, should not contain any significant amount of H2S.

The sphere has been in service for long and had been subjected to previous WFMT inspections. Itis assumed that most critical and evident cracking has been detected and repaired already. Crackingoccurred in the hardest regions of HAZ immediately adjacent to the weld cap edges and did notpropagate deeper than about 5 mm (0.2 in) because the steel was found to be softer at this depth. This istypical of multipass welding where subsequent beads temper previous beads below.

It was recommended that the sphere continued to be inspected and that the new cracks, if any,should be removed and repaired accordingly to reduce the probability of a failure. Temper bead methodwas mentioned as a possibility but PWHT was said to be the proven definitive remedy to this problem.The questions of whether to scrap the sphere, leave it as is, repair the cracks, PWHT the wholeequipment are always raised in these cases. The decision is up to the owner that needs to be advisedabout the risks involved. In principle the goal should be to perform WFMT inspection to 100% of thewelds, remove and repair all crack indications, perform PWHT to the whole sphere. The practice is thento reinspect the sphere after being put back in service to verify if any susceptibility remained. In thiscase, however, the likelihood of having further wet H2S cracking would be determined by the combinedprobability that there still are susceptible spots left and that there will be new H2S excursions. Theprobability of simultaneous occurrence of more than one event is much lower than the probability of

occurrence of a single event.

Stress-Assisted Corrosion in Steam Drum

We removed three samples to investigate the nature of the cracking at the blow down nozzle holeedge and surface of the bore of the one-inch nozzles of steam drum that operates at about 314°C(600°F). Inside surface cracking had been previously detected in this drum; however, this was the firstopportunity to remove samples for metallurgical analysis. A fourth sample was removed from a nearbylocation that contained typical inner diameter surface cracking. Prior to sample removal, white contrastwet magnetic particle examination had been performed to highlight the crack locations. The varioussamples were submitted to different metallurgical laboratories for analysis. This section covers theanalysis performed in only one sample removed from one-inch blowdown nozzle.

Figure 16 shows the small sample extracted in the one-inch blowdown nozzle. The front surface of the sample represents the inside surface of the steam drum. The back surface of the sample was the cutsurface produced by the hemispherical shaped saw cutting tool. Some cracks initiated on the holesurface rather than at the hole edge.

Figure 17 shows the metallographic cross section through the cracks, 0.5 mm (0.020 in) deep fromthe nozzle hole surface. The steam drum inside surface would be on the upper part of thephotomicrograph in Figure 17. There were three areas with cracks, though the cracked area in the middle

joined the cracking at the cut edge of the small sample. The cracks were not necessarily funnel-mouthed.They were opened and also had like internal caverns. Results of the metallurgical analysis showed thatthe cracks were transgranular, all wide opened, oxide filled. It is important to note that the cracksexhibited evidence that they were still active, not occurring from just a one-time event. No evidence of caustic stress corrosion cracking was found.

Most of the cracking was perpendicular to the inside surface of the steam drum, straight with somedegree of branching and with a few branches forming and growing evenly in the direction parallel to theplate, suggesting environmentally assisted cracking. The cracking observed was thus predominantly aform of environmentally assisted cracking that can best be described as corrosion fatigue or stressassisted corrosion17. It is also referred to as stress induced corrosion fatigue cracking, typically found inwater wetted boiler components loaded with pressure.

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The cracking mechanism involves breakage of the passive magnetite layer, initial penetration bylocalized corrosion, and growth by stress concentrating at the tip of these cracks and thermal cyclesassociated with start-ups and shutdowns. Because of the notch effect of existing cracks, they lose theirprotection and become exposed to oxygenated moisture during subsequent shutdowns. Acceleratedcorrosion may occur under these circumstances and form caverns at the tip of these cracks. As thecracks grow larger with time and further cycles occurred, these caverns are left behind, giving this verypeculiar crack appearance.

The design, age, material of construction and operating history needs to be reviewed in order toidentify the causes of the cracking problem and devise corrective measures. The generalrecommendations are to avoid operating with high frequency of start-up and shut down; minimize stressduring start-up by emphasizing gradual and steady start-ups; reduce or avoid pressure and temperaturefluctuations, sudden pressure lost, water hammer, and use a warm-up period as part of the start-upprocedure.

CONCLUSIONS

To ensure safe and reliable operations of the aging equipment for continued service a methodologyneeds to be developed implementing the existing codes. We have discussed evaluation of differentequipment using fitness for service approach along with metallurgical evaluation to assess of aging

equipment in refineries and chemical process plants. We identified the predominant degradationmechanism, and then evaluated the critical mechanical properties to ensure the mechanical integrity of the equipment. We further developed safe operating conditions for equipment as necessary.

In one case the required answer was to learn if the equipment could continue service for an extraperiod. This was done following the same approach but establishing inspection frequency to assure earlydetection of cracks or flaws and that they do not reach critical size. This prevents catastrophic failure.

There were also some evaluations that consisted more in finding and confirming the damagemechanism or providing an idea of the extent of damage. Cases of graphitization, wet H2S stresscorrosion and corrosion assisted cracking in carbon steel were described. The extraction of samplesallowed volumetric examination of the material.

In all the cases described, the evaluation was based on results that could be obtained in samplesextracted from the equipment. A technique was used and described to extract small samples out of which it was possible to machine sufficient specimens to perform different mechanical testing requiredfor the assessment.

ACKNOWLEDGMENTS

The authors acknowledge the contribution made by the owners and custodians of the equipmentsand companies where all these studies were performed and in providing the required information tosuccessfully complete each task. The patience, effort and contributions made by the reviewers are alsoappreciated.

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REFERENCES

1Fitness for Service: API Recommended Practice 579, 1

stEd., January 2000. API Publishing

Services, 1220 L Street, N.W., Washington, D.C. 20005.2

ASME Boiler and Pressure Vessel Code, Rules for Construction of Pressure Vessels, Section VIII,

Division 1, 2001 Edition, NY.3

API Standard 650, 10th

Edition, November 1998, Addendum 1, January 2000, Addendum 2,Novemebr 2001, entitled “Welded Steels Tanks for Oil Storage”, API Publishing Services, 1220 LStreet, NW, Washington, DC 20005.

4API Recommended Practice 934, 1st Ed., December 2000, entitled “Materials and FabricationRequirements for 2.25Cr-1Mo and 3Cr-1Mo Steel Heavy Wall Pressure Vessels for HighTemperature, High Pressure Hydrogen Service”, API Publishing Services, 1220 L Street, N.W.,Washington, D.C. 20005.

5 Tadao Iwadate: Pressurization Temperature of Pressure Vessels Made of Cr-Mo Steels, PVP-Vol288, Service Experience and Reliability Improvement: Nuclear, Fosil, and Petrochemical Plants,ASME 1994.

6 ASTM Standard E1290-02, “Standard Test Method for Crack-Tip Opening Displacement (CTOD)Fracture Toughness Measurement”, Philadelphia, PA, USA, ASTM International, 2002.

7Tahara, T.; Ishiguro, T., Effect of Short Temperature Excursion on Mechanical Properties of C-0.5Mo Steels, The Japan Steel Works, Ltd., Report No. R(PV) 92-027, November 24, 1992.

8R. O. Ritchie, “Near-threshold fatigue-crack propagation in steels”, International Metals Reviews,1979 Nos. 5 and 6, pp. 205-229.

9S. Suresh, C.M. Moss, and R.O. Ritchie: Proc. 2nd Japan Institute of Metals Int. Symp. onHydrogen in Metals", Sendal, Japan Nov. 1979, Paper 27B23.

10Broek, D., Elementary Engineering Fracture Mechanics, 4

threvised Edition (Martinus Nijhoff

Publishers, P. O. Box 163, 3300 AD Dordrecht, The Netherlands), 1986, pp. 260-266.11

Internal presentation, PDVSA-CITGO Annual Meeting on Materials and Reliability, Coke DrumCase, PDVSA-Intevep, Los Teques, Venezuela, Dec. 2001.

12API Recommended Practice 571, 1st Ed., December 2003, entitled “Damage Mechanisms AffectingFixed Equipment in the Refining Industry”, API Publishing Services, 1220 L Street, N.W.,Washington, D.C. 20005.

13Thielsch, H.; Phillips, E. M.; Jero, E. R.: Considerations in the evaluation of graphitization in pipingsystems. Presented in AWS National Fall Meeting, Chicago, IL, Nov. 1-5, 1954.

14 ASM Handbook, Formerly 9th Ed., Metals Hanbook, Vol. 11: Failure Analysis and Prevention,ASM International, Metals Park, OH, p.613, 1986.

15

J. E. Cantwell: LPG Storage Vessel Cracking Experience. NACE CORROSION 1988, paper #157(Houston, Texas: NACE International, 1988).

16Marvin Mehler: Serviceability Analysis of LPG Sphere with Hydrogen Sulfide Cracking. NACECORROSION 1990, paper #205 (Houston, Texas: NACE International, 1990).

17P. B. Desch, J. J. Dillon, and S.H.M. Vrijhoeven: Case histories of stress assisted corrosion inboilers. NACE CORROSION 2004, paper #04516 (Houston, Texas: NACE International, 2004).

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Table 1. Scoop sample and divot dimensions

Specim

en No.

Max. Useable

Thickness of Specimen,

mm (inches)

Max. Depth of

Divot, mm (inches)

Diameter of Divot In

Long. Direction, mm

(inches)

Divot Diameter in

Circum. Direction,

mm (inches)

5A 12.0 (0.472) 15.1 (0.595) 54 (2.142) 52 (2.033)

5B 12.2 (0.480) 14.9 (0.588) 54 (2.114) 53 (2.087)

5C 11.8 (0.466) 14.5 (0.570) 53 (2.102) 53 (2.085)

6A 12.0 (0.474) 14.0 (0.550) 53 (2.089) 53 (2.072)

6B 11.8 (0.465) 14.4 (0.566) 53 (2.108) 53 (2.086)

6C 11.7 (0.459) 13.8 (0.542) 53 (2.092) 52 (2.062)

Table 2. CTOD test results using specimens of Coke Drums

Specimen Drum / Location conditionKmax

(ksi√in)JIntegral (lb/in)

CTOD

(mm)

CTOD

(inches)

A1 - 1of 2 Coke Drum 1 / Bulge Tail 57.0 303.0 0.0980 3.806 x 10-3

A2 - 2of 2 Coke Drum 1 / Bulge Crown 55.7 144.6 0.1143 4.500 x 10-3

B1 - 1of 2 Coke Drum 2 / Bulge Tail 40.1 48.9 0.0253 0.996 x 10-3

B2 – 1of 2 Coke Drum 2 / Bulge Crown 53.6 199.7 0.1222 4.810 x 10-3

B2 – 2of 2 Coke Drum 2/ Bulge Crown 52.6 111.8 0.1232 4.850 x 10-3

B3 - 1of 1 Coke Drum 2 / No bulged 53.1 203.0 0.1329 5.230 x 10-3

Table 3. Mechanical testing of specimens machined out of scoop samples

Lab Identification Sample 1 Sample 2

Dimensions 5.3x3.25 mm (0.209x0.128 inch) 5.6x3.3 mm (0.220x0.130 inch)

Area 17 mm2 (0.0267 inch2) 18 mm2 (0.0286 inch2)

Yield Strength 313 MPa (45,400 psig) 296 MPa (42,900 psig)

Tensile Strength 434 MPa (63,000 psig) 409 MPa (59,300 psig)Elongation 15.1% 11.7%

Reduction in Area 20.6% 21.0%

Fracture Appearance Ductile Ductile

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Figure 1. As-removed scoop samples

0

10

20

30

40

50

60

70

80

0 5 10 15 20 25 30 35

Temperature, °F

C h a r p y I m p a c t E n e r g y , f t - l b

Tank 6

Tank 5

Figure 2. Charpy impact energy values for storage tanks

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Figure 3. Divot after removal of a scoop sample in reactor top head

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Ring #2Base Metal

Top HeadBase Metal

6050403020100-10-20

0

20

40

60

80

100

Ring #1Base Metal

Ring #1Weld Metal

Temperature, °F

C h a r p y I m p a c t E n e r g y ,

f t - l b

Figure 4. Charpy V-Notch test results from reactor specimens

10

20

30

40

50

-10 0 10 20 30 40 50 60

Regression Plot95% Confidence Bands

23 ft-lb

A b s o r b e d E n e r g y

, f t - l b

Temperature, °F

Figure 5. Best fit line and associated 95% confidence curves for Charpy energyvalues

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400

600

800

1000

1200

1400

1600

1800

-70 -50 -30 -10 10 30 50

Temperature, °F

P r e s s u r e , p s i g

Figure 6. Safe envelope to avoid brittle fracture in the reactor

0.0

0.5

1.0

1.5

2.0

0 5 10 15 20 25 30 35 40

Crack Length (inches)

C r a c k D e p t h ( i n c

h e s )

unacceptable

acceptablereference crack

Figure 7. Damage tolerance curve assuming cracks parallel to longitudinal welds,120 kg/cm2 (1700 psig) at 6°C (43°F)

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0

1

2

3

4

5

6

7

8

9

10

-80 -60 -40 -20 0 20 40 60

Presurization Temperature, °F

C r a c k D i m e n s i o n s

, i n c h e s Crack Length, 2c

Crack Depth, a

Figure 8. Fracture tolerance signature with cracks parallel to longitudinal welds,using 48 kg/cm2 (680 psig) to 120 kg/cm2 (1700 psig)

Figure 9. Three-point bend specimens for Crack Tip Opening Displacement(CTOD)6 tests

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0.25

0.35

0.45

0.55

0.65

0.75

0 20 40 60 80 100 120 140 160 180 200 220 240

Cycles

C r a c k d e p t h " a " ,

i n c h e s Through Wall Crack

Cyclic Stress 43 ksi

Cyclic Stress 32.6 Ksi

da/dN = 2.35 x 10-9

K3.64

(a)

1.5

1.6

1.7

1.8

1.9

2.0

2.1

2.2

2.3

0 20 40 60 80 100 120 140 160 180 200 220 240

Cycles

C r a c k L e n g t h " 2 c " , i n c h e ) Through Wall Crack

Cyclic Stress 43 ksi

Cyclic Stress 32.6 Ksi

da/dN = 2.35 x 10-9

K3.64

(b)

Figure 10. Fatigue crack length growth, cyclic stress (a) 225 MPa; (b) 296 MPa(min yield strength)

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Figure 11. Crack formed during bending test of graphitized steel. Scale in tenthsof an inch

Figure 12. Random graphitization in base metal and aligned in the metal adjacentto HAZ. 5% Nital etch

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Figure 13. Graphite nodules in a pearlite-free ferritic matrix. 5% NItal etch, 80X

Figure 14. Metallographic section through scoop sample

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Figure 15. Microhardness HV1 survey done on scoop sample

Figure 16. Scoop sample taken from a one-inch nozzle

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Figure 17. Metallographic cross section through the hole cracks. Etched in 5%Nital, 14X

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