Effect of Macrostructure onCatalyst Tube Damage and

Creep Rupture Properties of HK40

Susceptibility to creep damages is mainly affected by themacrostructure of the metal. Therefore, the macrostructure of thespeciman should be inspected and recorded for discussion of thecreep-rupture properties of heat resisting alloy castings.

Takao Kawai, Katsuaki TakemuraToshikazu Shibasaki, and Takaaki Mohri

Chiyoda Chemical Engineering & Construction Co., Ltd.Yokohama, Japan

Catalyst tubes of steam reformers are usedunder conditions of high temperatures andhigh pressures, consequently, they sufferfrom creep damages during long term opera-tion.

The authors have investigated the prog-ress of creep damage on catalyst tubes inreformer furnaces of Hydrogen and AmmoniaPlants constructed by Chiyoda Chemical Engi-neering & Construction Co., Ltd. to predictthe remaining tube life for deciding when theyhave to be replaced and to check the changein metal structure and mechanical propertiescaused by ageing during long term operation.For the investigation, one to several tubeswere extracted from each furnace as represen-tative samples of the tubes. Until the end of1977, after at least three years of service, atotal of 53 unfailing tubes had been removedfrom a total of 15 furnaces 34 times.

In the past, some reports have said thatmost of creep failures of catalyst tubes weredue to overheating. However, recently, it hasbeen increasingly recognized that even tubesunder normal operating conditions withoutoverheating have developed creep cracks muchearlier than the expected rupture time due tothe existence of high thermal stresses causedby temperature difference in the tube wall. Inthe same manner, creep fissures or creepvoids were often observed in the destructiveinvestigations we made, although apparentlythere were no tubes which had suffered fromoverheating during services. It was inducedfrom further observations that the rate of0149-3701-80-3938 $01.00 © 1980American Institute of Chemical Engineers

creep damage, in the base metal taking theform of longitudinal cracks or in the weld ascircumferential cracking (which accounts forthe greater part of the total weld cracks), ismuch different even though the tube had beenexposed to the same service conditions.

On the other hand, a computer programon creep analysis using finite element methodswas developed to estimate the remaining life ofused catalyst tubes. However, the greatdifference between the calculated results andthe actual metallographic appearance by de-structive essays suggested that this problem ismuch more complicated than imagined.

Accordingly, the authors began a moredetailed study to assess the available data oncreep rupture properties of heat resisting castalloys, to inspect metallographic behavior oftubes after actual service, and to conductcreep-rupture tests on different kinds of metalstructures.

CREEP-RUPTURE STRENGTH FOR SPUN CAST

HK40 TUBES

Since 1964, Estruch and Lyth's Larson-Miller diagram with the Larson-Miller constantC=15 has influenced to decide the designstress for HK40 tubes. (1) Untill now no otherpractical paper has been published on thismatter, although in recent years a certainamount of new data has been taken whichindicates that the long term creep-rupturestrength is somewhat lower than Estruch'sCurve.

In 1974 the National Research Institute

119

Tsbte 1 Summary of Croap-Rupture Teils Data by NRIM

( The Nation«) Research Instituts for Matali ) (2)

Tut Condition! Craep-Ruptura Times

Temperature Stress

°C N/mm1

800 98.0

68.6

52.9

33.3

900 52.9

33.3

26.5

19.6

1,000 26.5

19.6

13.7

9.8

1,050 9.8

1,100 9.8

Number

of

Data

Points

14

14

14

12

14

14

14

14

11

14

14

14

14

14

Longest

Hours

264.8

1,274.6

3,053.8

25.418.2

446.6

2,406.4

5,198.1

20,874.6

764.5

1,295.3

4,324.2

16,779.7

4,871.3

1,370.2

Shortest

Hours

28.8

365.8

839.0

7,804.8

69.6

600.2

1,228.8

4,686.5

185.3

442.4

1,391.8

3,191.8

1,088.5

240.7

Arithmotic

Masn

Hours

122.5

716.4

1,952.0

14,878.2

174.6

1,050.4

2,674.1

12,113.3

283.1

808.6

2,821.5

10,258.7

2,694.8

577.8

Logarithmic

Mem

Hours

103.3

672.6

1,852.1

14,024.0

155.4

977.4

2.454.9

10.933.9

258.3

776.7

2,685.3

9,285.5

2,483.6

518.5

for Metals (NRIM) of Japan issued interim datasheets on creep-rupture tests for spun castHK40 tubes. (2) These tests were carriedout on 14 spun cast tubes which had beensupplied by Japanese foundries in 1967. Thetotal number of test points were 194, extendingover the range 800-1,100°C and including testsof over 25,000 hours. A summary of thesetest results is shown in Table 1.

Based on the NRIM data, logarithmic meanrupture times were plotted on Estruch's dia-gram as shown in Figure 1. It was found thatfor all test temperatures, the longer therupture times are, the lower the rupturestresses are compared to Estruch's values.This tendency was evidently the same on thedata of all individual sample tubes. It wasrecognized that the Larson-Miller constant C=15taken by Estruch was not suitable since theconstant should give the least dispersion aboutthe central tendency line. Namely, the indi-vidual temperature lines should make thesmallest angle with the master line by determin-ing a suitable Larson-Miller constant.

Based on NRIM logarithmic mean livesexcept for the data from 1,050°C and 1,100°C,the least square method was applied to deter-mine the value of the Larson-Miller constanttogether with constants of the equation relatingthe Larson-Miller parameter, P, to the rupturestress. The best fitting curve (hereinaftercalled the KS curve) came from the followingequation. (3) (4)

P = 16.955 - 4.174 X - 0.519 X2

200

100

- 50

"Ez

20

10

5 —

19 20 21 22

P * T (log t + 15.0) x Iff'

23 24

( X Hour«)

Figure 1. Logarithmic mean rupture times of NRIM dataplotted on Estruch's Larson-Millr diagram.

where, p = T(log t + 9.4) x 103

X = logo 2a = Stress in Kg/mmT = Temperature in °Kt = Rupture time in hours

The KS curve in the Larson-Miller diagramwith the constant C=9.4 is illustrated in Figure2 and in which the logarithmic mean lives ofNRIM data, which are the same as those plottedin Figure 1, are less dispersed from the centraltendency line of the K S curve.

Figure 3 gives a comparison of averagerupture stresses for a life of 100,000 hourscalculated from the KS equation and Estruch's

200

100

so

20

10

IO 800'CA 900'C

D 1.000 "C

X^xj Xi

12 13 14 15f - T (log t + 9.4) X

16 17' ( °K. Hours )

18

Figure 2. The KS curve for spun cast HK40 materials withthe Larson-Miller constant c=9.4.

120

100

so

20

E

Z

10

O Hot» From Log(Hr)

— Logic) Cum

ASTM A567

800 900Tamper« tuts

1.000

Figure 3. Comparison of average rupture stresses in100,00 hr. for HK40 materials.

equation and the expectancy shown in ASTMA567. It is evident that the KS curve isabout 35% lower than the values of Estruch andASTM. These big differences can be explainedby the selection of the Larson-Miller constant.

It is necessary at this point to discussthe errors caused by the handling of the N RIMdata. Though the logarithmic mean life wasused for the calculation of the K S curve as arepresentative of 14 individual test points, theresults were compared with those obtainedusing the average life or all individual lives,and the differences were within a few percent.The KS curve excludes the data from tests at1,050°C, and 1,100°C but in the cases includ-ing these data, the Larson-Miller constant wascalculated as C=10.0. This value was somewhathigher than C=9.4 of the KS curve; therefore,the estimated mean rupture stresses in 100,000hours become 5 to 10% higher than the K Scurve. However, considering the effect ofsignificant metallurgical changes on rupturestrength at those higher temperatures, data of1,050°C and 1,100°C should not be evaluated.

On the other hand, the log-log methodfor estimating mean rupture stress in 100,000hours was also applied to the N RIM data fortemperatures of 800°C, 900°C and 1,000°C.These are plotted in Figure 3 and show thatthe values are in reasonable agreement withthe K S curve.

Thus the conclusion reached was that themean rupture stresses for spun cast HK40material are expressed more reasonably by theKS curve as far as they are based on theN RIM data.

Next we attempted to determine the most

significant factor affecting creep-rupturestrength of HK40 tubes. It is generally believ-ed that the spun cast product is stronger thanthe static cast product since the quality of thespun cast material is better in uniformity,purity, density, etc. However, the creep-rupture strength estimated from the K S equa-tion which represents 163 data points obtainedexclusively on spun cast products is signifi-cantly lower than other published values,although some include data of static castproducts. The investigation was continued tofind out what is most attributable to creep-rupture strength.

First, the relationship between the tensilestrength at room temperature and the rupturetime was inspected. And the results was thesame as it is said, i.e., for short term rupturetests of around some hundred hours a correla-tion was observed, however for long termrupture tests of more than 10,000 hours, thetensile strength at room temperature can notaffect them any more.

The same relationship was recognizedbetween the elevated temperature tensilestrength and the rupture time. It is generallysaid that the tensile property does not have adependable relationship to the long term creep-rupture property, but sometimes the fact thatthe long term creep-rupture strength is alsoindependent from the short term rupture testas far as the tensile strength is concerned isoverlooked. This non-relationship was alsoindicated from N RIM data.

In this regards, the effect of the durationof the creep-rupture test should be checked tounderstand the evolution of the problem.Based on rupture times of NRIM data, the

60

160I

$40

I

L20 -

AO

O

A

A 800 °CO 900 °CV 1,000 °c

o

A V

1,050 °C

1,100°C

A

i100 200 500 1,000 2,000 5,000

Average Rupture Time { Hours )10.000

Figure 4. Standard deviation ratio vs. average rupturetimes.

121

Standard deviation ratio for each test conditionwas inspected, where the standard deviationratio is defined as the value of the standarddeviation of rupture times divided by averagerupture time. This result is plotted in Figure4 and explains that for cases where the creep-rupture test duration is less than about 600hours, the standard deviation ratio is evidentlyhigher than the values for longer test durationcases, independent of the test temperature.This makes us think that the error mayincrease more and more for estimation of longterm rupture times when the percentage of thecreep-rupture data points from shorter testdurations are increased. Therefore, reasonablecare should be paid to the treatment of individ-ual creep-rupture test data.

Figure 5 shows the relationship betweenthe carbon contents of HK40 spun cast tubeand creep-rupture time. The data for shortterm rupture (900°C, 53 N/mm 2 and 1,100°C,9.8 N/mm2 ) are evidently affected by carboncontent. The rupture time increases in pro-portion to the carbon level. However, for longterm data around 12,000 hours (900°C, 19.6N / m m 2 ) carbon content is not attributable tothe rupture time. Consequently, the carbonlevel within the HK40 range has no pronounced

20,000

10,000

5,000

eg5 '2,000

oJE 1,000

£

t 500cc

200

100

50

V

vv

V

v v

1,100 "C

Q 9.8 N/mm3

" ° 8°o

AA ^A

A—

i i i i i i0.35

^2 szv V v ̂900 °C

19.6 N/mm1

0°

°§ 2O W tA

A

A A

7S ̂ »GO °c52.9 N/mm1

1 1 1 I 1 I 1

0.40 0.45Carbon Content (%)

Figure 5. Effect of carbon level on creep-rupture times forspun cast HK40 tubes.

influence on long term creep-rupture times asexpected for industrial service. However, itis still widely believed that a higher carbonratio increases creep-rupture time over longterm service. Actually a high carbon contentmay contribute only to lessening the ductilityof the alloy rather than increasing the tubelife.

Incidentally, the effects of other elementsof the HK40 alloy on rupture time were inspect-ed and no evident relationship was found.

Recently, some reports said that thesigma phase shortens creep-rupture time andthat for HK40 material the formation of thesigma phase may occur in long term servicewhen the average electron-vacancy number Nvis higher than 2.70, and the volume of theprecipitated sigma phase may increase inproportion to the average electron-vacancynumber. (5>) For tests at 800°C and 33.3N/mm 2 , the relatipnship between the creep-rupture time and the average electron-vacancynumber was examined. The rupture timeswere somewhat shortened in accordance withthe increase of the average electron-vacancynumber, namely the increase of the sigmaphase. However, considering the scatteredrange of the individual data, this tendencywas not as strong as had been reported. Inhigher temperature cases, for example at900°C, the average electron-vacancy numberhad no correlation to the rupture time, sincethe temperature had been within a high rangewhere no sigma phase may be precipitated.

Thus many items have been examined tofind the most attributable reason for creep-rupture behavior, however the answer did notappear from the chemical compositions ormechanical properties. We deduced from theseresults that the main reason was still to befound. It was finally found in the macrostruc-ture as will be discussed later.

CREEP DAMAGE OF SPUN CAST TUBES

During the metallurgical investigations oncatalyst tubes after long service, severallongitudinal creep cracks were observed in thebase metal of the catalyst tubes, even thoughthe tubes were used under uniform and goodtemperature control without local overheating.Also, there were many types of crack propaga-tion caused by creep damage in the tube.

Photograph 1 is an example of severelongitudinal creep damage in the base metal.It seems as if the tube ruptured but no leak-age was detected during service. There is nodifference in severity of creep damage in thiscross section. The implication is that the tubehad been under uniform temperature distribu-tion though one side of the tube was facing

122

Photograph I. Severe creep damage around thecircumference of ara HK4© catalysttube section.

the burner wail.In Photograph 2 (a) the creep damage is

indicated near the inner surface and is devel-oped continuously along the columnar grainboundaries from the inside to outside surface.The appearance of creep damage changesgradually from fissures to aligned voids toisolated voids. This propagation type of creepdamages agrees closely to the results calculatedby computer analysis. Photograph 2(b) showsthe macrostructure of the tube having a fullycolumnar grain structure.

Photograph 3 is a sample taken from thesame catalyst tube shown in Photograph 1.However, in Photograph 3 (a) no crack wasobserved in the outer tube wall, althoughisolated creep voids were found by microscopicexamination. Photograph 3(b) shows the macro-structure of a tube section which exhibits acombination of columnar grains for the insideand equiaxed grains for the outside. Thedevelopment of the creep cracks seemed tostop at the boundary of the columnar structureand the equiaxed structure.

Photograph 4 is another sample tube usedunder the same service conditions as the tubeshown in Photograph 1. The macrostructureshown in Photograph 4(b) is unusual for aspun cast tube. Each quarter layer of thetube wall has a different macrostructure,namely from the outside fine equiaxed grains,fine columnar grains, fine equiaxed grains,

Photograph 2. Advanced creep damage of a fully columnarHK40 tube section.

Photograph 3. Creep damage of an HK40 tube section withfine equiaxed grains and columnar grains.

Photograph 4. Creep damage of an unusual macrostruc-ture HK4® tube section.

and coarse equiaxed grains. As shown inPhotograph 4 (a) creep cracks were found inlayers of coarse equiaxed grains and columnargrains. However, only creep voids wereobserved in the inner fine equiaxed grains

123

Photograph 5. Creep damage in longitudinal section withdifferent macrostructures.

layer which was located between both crackedlayers • The propagation of creep damageseemed to be delayed in this layer since it waslocated in the inner portion rather than thecracked layer of columnar grains. This typeof damage propagation is one of types whichcannot be predicted by computer calculation.

Photograph 5 shows another sample ofcreep damage, which exhibits a longitudinalcross section. Though there were no differ-ences in chemical composition between bothcast tubes, and although they had been usedin the same operating conditions, the appear-ance of creep damage was quite differentbetween the coarse columnar structure tubeand the fine equiaxed structure tube. Creepcracks were found only in the tube of columnarstructure. /

From these investigations, it was confirm-ed that susceptibility to creep damage of thecatalyst tube is influenced by the macrostruc-ture of the metal.

CREEP DAMAGE IN WELD METAL

In some catalyst tubes circumferentialcracks were observed in the middle of thegirth weld as shown in Photograph 5. It isevident that the cracks are creep damagebecause many creep voids are found at the endof the cracks and in the vicinity of the cracks.The circumferential cracks in the weld mayhave been the result of thermal-gradientstresses since the stress due to inner pressureacts in the longitudianl direction is only halfof the level of the stress in the circumfsrentialdirection (hoop stress). In our investigationsa total of 53 tubes were taken 34 times from atotal of 15 furnaces. The circumferentialcreep cracks were found 20 times from 11furnaces. The locations of the cracked girthwelds were not fixed, however in some fur-

Photograph 6. No creep damage in the weld metal, thoughin base metals severe damages are evident.

naces, they were limited to the top portion orbottom portion.

The furnaces had been designed withnearly the same conditions and had the sameoperation manuals, however, in several fur-naces no circumferential cracks were foundeven though the base metal had been damagedseverely. Photograph 6(a) shows a samplewhere a few creep voids were found in theweld, although in both base metals the longi-tudinal creep cracks extended almost throughthe tube wall.

Comparing the macrostructure of the weldbetween Photograph 5(b) and Photograph 6(b),the difference in the grain growth of the metalstructure can be recognized easily. Thedetailed macrostructure of welds shown inPhotographs 5 and 6 were inspected from threedifferent views.

The detailed macrostructures of a crackedweld are shown in Photograph 7. Columnargrains of the weld metal are developed continu-ously and the direction of the grain growth isperpendicular to the outside surface, namely,in a radial direction. This type of macrostruc-ture in the weld in called Type P. (6)

Photograph 8 shows the detailed macro-structures of a noncracked weld. Columnargrains are inclined and twisted both to thelongitudinal and to the circumferential direc-tions of the tube. This type of macrostructurein the weld is called type T.^6)

The big difference of the macrostructurein the weld between Type P and Type T maybe caused by a difference in welding proce-dures. It was notable that no circumferentialcracks were found in the Type T weld in ourinvestigations.

124

Photograph 7. Macrostructures of Type-P weld in three-different sections.

CREEP TEST ON A THICK-WALL S-PUN-CAST

HK40 TUBE

In order to ascertain the effects of macro-structure on creep properties, a very thick-wall spun-cast HK40 tube was made and creeptests were conducted. The outer diameter ofthe sample tube was 293 mm and the wallthickness was 98 mm. There was no differencein chemical composition in the outer, middle orinner portions of the tube wall.

The macrostructure of this sample tubeconsists of three main layers» that is equiaxeugrains, columnar grains and again equiaxedgrains from the outside. As shown in Photo-graph 9, the following seven kinds of speci-mens were taken from each layer in differentdirections.

Specimens A were taken logitudinally from thecolumnar structure.They have a perpendicular stress axis tothe direction of the grain growth.

Specimens B were taken circumferentially fromthe columnar structure.

Photograph 8. Macrostructures of Type-T weld in threedifferent sections.

«- I , ! ' l i' l l -.5 6 7 8 9

,!,|„I,„>, I ) l.,.'.| I II

,!„f,,,.( ..l, , , ).{!., M, ,, ,1. ! .,« ),!., (>Il, j I , , j

S 6 7 S 9 I J O l 2 3 d 5 €,i-.hi,l TM',,!, 'i. kiMLai :I,i!l'j!!j..!.,'A,J.",l l > l . l « , l

Photograph 9. Macrostructure and specimens in a thick-wail spun-cast HK40 tube.

They also have a perpendicular stressaxis to the direction of the grain growth.

125

Etatisation

Location and Dirvctioft of Speeimem

Figure 6. Effect of macrostructures on room temp, tensileproperties for a thick-wall HK40 tube.

100 300 1000 3000

Rupture Tim« ( Mrs )

* B C D E F G

Location and Direction of Specimens

Figure 7. Effect of macrostructures on creep-rupturestrength for a thick-wall HK40 tube.

Specimens C were takencolumnar structure.They have a paralleldirection of the grain

Specimens D and F wereand circumferentiallyequiaxed grain layer,

Specimens E and G wereand circumferentiallyequiaxed grain layer,

radially from the

stress axis to thegrowth,taken longitudinally

from the outerrespectively,taken longitudinally

from the innerrespectively.

Figure 6 shows a comparison of the tensileproperties at room temperature for each of thespecimens. Specimens C have extremely hightensile strength and high elongation comparedwith those of other specimens.

As to the tensile properties at elevatedtemperature, the ultimate tensile strength andthe yeild strength were almost the same valuefor all specimens. However, the elongation ofspecimens C showed an extremely higher valuethan the others.

The creep-rupture tests at 850°C, 900°Cand 950°C were conducted as shown in Figure7 and the creep-rupture elongations are plottedin Figure 8. Comparing the effect of the

30 100 300 1000 3000

Rupture Time ! Mrs )

A B C DLocation anu Direction of Specimens

Figure 8. Effect of macrostructures on creep-ruptureelongation for a thick-wall HK40 tube.

126

stress direction to the grain growth of thecolumnar structure on the creep-rupturestrength and on the creep-rupture elongation,specimens C have extremely higher values thanspecimens A and B. Namely, when the speci-mens are taken from the columnar structureand are stressed parallel to the direction ofthe grain growth, they have the highestcreep-rupture strength and the best ductility.

For specimens D, E, F, and G, nodifference in creep-rupture properties wasfound. This means that for the equiaxedgrains, the stress direction and the location ofthe specimens in the thick tube wall do notaffect the creep-rupture properties, althoughthe grain size of specimens D and F whichwere taken from the outer equiaxed layer wassimilar to that of specimens E and G takenfrom the inner equiaxed layer.

Both the creep-rupture strength and thecreep-rupture elongation of equiaxed struc-tures, for specimens D, E, F, and G, showmedium values between those of specimens Cand specimens A and B. That is, creep-rupture properties of the equiaxed structureare better than those of the columnar structureas long as the stress acts perpendicularly tothe direction of the columnar grain growth.

When the creep curves of each specimenwere compared, it was interesting to note thatthe secondary creep rates of each specimenwere essentially the same; therefore, the bigdifference in the rupture times between themwere caused by the difference in the durationof the secondary stage of the creep curves.

Macro Etched 1N KOH Etched

The difference of the creep-rupture elonga-tions were mainly seen in the difference of thestrains in the tertiary creep stage.

EFFECT OF MACROSTRUCTURE ON CREEP

PROPERTIES

Although specimens A, B and C are takenfrom the same columnar grains layer, specimensC show entirely different properties especiallyin creep characteristics.

Photograph 10 shows a fractured specimenA, sectioned parallel to the stress axis throughthe approximate centre of the specimen. Thefracture plane is rather flat and faces thestress direction, and a few sharp and longcracks are observed near the ruptured surface.Cracks are fairly straight and perpendicular tothe stress axis and are observed preferentiallyalong the grain boundaries which continuestraight and confront the stress direction inspecimen A.

Photograph 11 shows a section of a frac-tured specimen C. The ruptured surface hasan irregular shape. Numerous small dividedcreep cracks are observed in the specimen,and some of them are elongated in the stressdirection.

Photographs 10 and 11 show also micro-photographs of sectioned specimens A and C,respectively. Creep cracks and creep voidsare seen at the grain boundaries, namely ateutectic carbide and austenite interfaces. Asshown in the Photographs, the appearance ofthe creep voids are not of the wedge type andthey are formed and propagated preferencially

\

Macro Etched IN KOH Etched

Photograph 10. Axial section of a fractured Specimen A,tested at 850°C and 59.0 N/mm2.

Photograph 11. Axial section of a fractured Speelman C,tested at 850°C and 59.0 N/mm2.

127

where grain bundaries are mainly perpendicularto the stress axis.

In specimens A and B, since the grainboundaries of the columnar structure continuein a plane and confront the stress axis, it isconsidered that creep cracks propagate easilyalong the grain boundaries. On the otherhand, in the case of specimens C, creepcracks can only develop with a large deforma-tion because small cracks are hardly linkedwith each other since the grain boundaries arenot aligned continuously in a perpendicularplane against the stress axis. This may also beimplied from the creep curves. They show thesame creep rates at the secondary stage,however the initiation of the tertiary creep ofspecimens C begins rather later than that ofspecimens A and B. For specimens C, muchstrain is required to propagate the creepcracks. This is considered to be the reasonwhy the rupture ductilities and strengths ofspecimens C are higher than those of specimensA anb B.

For specimens D, E, F and G which weretaken from the equiaxed grain layers, theshape and distribution of creep cracks aremidway between those of specimens A or B andspecimens C, as shown in Photograph 12. Thecontinuity of grain boundaries which confrontthe stress axis is between those of specimensA or B and specimens C. For this reason,creep properties such as the initiation of thetertiary creep, rupture strengths and ruptureelongations show medium values. However, itis easily recongnized that when the equiaxedgrains coarsen, the creep properties will come

As Polished Macro Etched IN KOH Etched

Photograph 12. Axial section of a fractured Specimen D,tested at 850°C and 59.0 N/mm2.

close to those of specimens A and B. In ourcreep tests, there were no big differencesbetween the grain size of specimens D and Fagainst those of specimens E and G, althoughthey were taken from the outer equiaxed layerand the inner equiaxed layer respectively.

DISCUSSIONS

The results obtained from creep-rupturetests on the thick-wall spun-cast HK40 tubecan clearly explain the main cause of thecomplicated problems which we faced in theprevious stage.

In regard to the creep-rupture strengthof spun cast HK40 material, it was confirmedthat the macrostructure of the metal is themost significant factor affecting rupture time.The K S curve represents data obtained exclu-sively on spun cast products. Though it hadnot been inspected by NRIM, the macrostruc-tures of the greater part of the NRIM speci-mens were most probably of columnar struc-tures; therefore, grain boundaries wereprobably perpendicular to the stress axiscorresponding to specimens A and B whichwere the weakest in our creep-rupture tests.On the contrary, in regard to the macrostruc-ture and the direction of grain boundaries tothe stress axis of specimens for static castproducts, they are not defined since proce-dures and methods of taking specimens are notstandardized.

Thus, the long term creep-rupturestresses calculated by the KS equation arerather lower than other published values suchas ASTM A567 which includes data from staticcast products. It was also understood that themain reason for the wide spread of creep-rupture data for the same test conditons andfor the same alloy material is the difference inthe macrostructure and stress direction to thegrain boundaries of the test specimens. Inany case, the results obtained through ourmetallographic investigations show that themacrostructure data of the test specimens isindispensable in the evaluation of creep-rupture data.

As to the difference in the rate of longi-tudinal creep damage of the catalyst tube, thereasons can be explained as follows. Theinside portion of the catalyst tube wall suffersfrom high thermal stresses and inner pressurestresses at high temperature. Therefore, inaddition to the rupture strength, higher creeprupture elongation are required for the catalysttube in order to reduce the thermal stresses toa lower level. Against these arduous serviceconditions the fine equiaxed grain structurehas better characteristics, as confirmed byspecimens D, E, F and G in creep-rupture

128

test, higher creep-rupture strength and betterductility than those of columnar grain structurewhenever stress works perpendicular to thedirection of columnar grain boundaries.Because the direction of grain growth for thecolumnar structure of the spun cast tube ismainly in a radial direction, the hoop stressand longitudinal stress work perpendicularly tothe grain boundaries of the columnar structuresas in specimens A and B. Coarse equiaxedstructures also have large grain boundarieswhich may often face the stress directions.Consequently, the coarse structures cannotprevent the propagation of creep damage. Forthis reason, as experimentally shown, thepropagation of creep damage in the catalysttube is related to the macrostructure of themetal.

Similarly the reason for different sensitiv-ities to circumferential creep cracks betweenwelds of Type P and Type T is due to thedifference of the creep-rupture propertiesbetween them. In the case of Type P, grainboundaries of the columnar structure in theweld are aligned perpendicularly to the longitu-dianl stresses and also to the circumferenctialstresses. The same relationship between thestress axis and grain boundary holds forspecimens A and B which were the weakestand were less elongated on the creep-rupturetests. On the contrary, grain boundaries forType T welds are inclined or twisted to bothlongitudinal and circumferential stress direc-tions | therefore, at least some parts of thecolumnar structure have a portion parallel tothe stress axis. The stress condition of thisparallel portion corresponds to specimens C inthe creep-rupture tests. When stress wasapplied parallel to the direction of the grainboundaries, the columnar structure showedhigher creep ductility than when stress wasapplied perpendicularly to the grain boundari-es. This indicates that weld metal which has aType T macrostructure could relax the thermalstress to a lower level than that of Type P.Furthermore, the creep-rupture strength ofweld metal of Type T macrostructures isexpected to be higher than that of Type P.Therefore Type T welds should have moreresistance to circumferential creep damage thanType P welds.

Photograph 13. Desirable macrostructure of spun-casttubes for reformer service.

CONCLUSION

Thus we reached the conclusion thatsusceptibility to creep damages is mainlyaffected by the macrostructure of the metal.

The following suggestions offer the mostdesirable macrostructure for steam reformercatalyst tubes.

(1) For spun cast tubes, a certain portion ofthe inner tube wall should have fineequiaxed grains where thermal stressesand inner pressure stresses are at a highlevel, and the remaining outer portion ofthe wall should be of a fine columnarstructure in order to confirm the sound-ness of the casting. One preferableexample is shown in Photograph 13.

(2) As to the girth weld, a Type T macro-structure is desired where the graingrowth of the columnar structure isinclined and twisted to the hoop andlongitudinal stress axes. In order to geta Type T macrostructure the weldingprocedure and the shape of the bevelledends should be selected carefully.

It was also found that the macrostructureof the specimen should be inspected andrecorded for discussion of the creep-ruptureproperties of heat resisting alloy castings.

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LITERATURE CITED

1. B. Estruch and C. Lyth, "Materials Tech-nology in Steam Reforming Processes,"p.29, Pergamon Press Ltd., London (1966)

2. National Research Institute for Metals,"NRIM Creep Data Sheet, No. 16," (1974)

3. T. Kawai and T. Mohri, The llth AutumnMeeting of Japan Chemical EngineeringSociety, 651 (1977)

4. T. Kawai and K. Takemura, The 20thSymposium of Japan Corrosion EngineeringSociety, 13 (1978)

5. S. Ohta, Report of the 123 Committee onHeat-Resisting Metals and Alloys, JapanSociety for the Promotion of Science, 18,383 (1977)

6. K. Takemura, K. Naitoh and T. Shibasaki,The Third International Symposium ofJapan Welding Society, 373 (1978)

KAWAI, T. TAKEMURA, K. SHIBASAKI, T. MOHRI, T.

DISCUSSION

ROBERT PRESCOTT, C.F.BRÂUN: When you established thestress to produce rupture in a 100,000 hr., what minimum carboncontent were you considering?

T. KAWAI, Chiyoda: This KS curve is based on spun-cast HK 40tubes, thus within the carbon range of HK 40, meaning 0.35 to

0.45%.

PRESCOTT: We are finding consistently that the initial failuresin reformer furnaces are the lower carbon tubes. Therefore,many people have increased the minimum carbon content up to0.40 or 0.42%.

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