Engineering Failure Analysis 45 (2014) 456–469

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Engineering Failure Analysis

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Evaluation of creep damage of INCOLOY 800HT pigtailsin a refinery steam reformer unit

http://dx.doi.org/10.1016/j.engfailanal.2014.07.0171350-6307/� 2014 Elsevier Ltd. All rights reserved.

⇑ Corresponding author.E-mail addresses: lspyrou@ireteth.certh.gr (L.A. Spyrou), pasarafoglou@uth.gr (P.I. Sarafoglou), aravas@uth.gr (N. Aravas), hgreg@m

(G.N. Haidemenopoulos).

L.A. Spyrou a,⇑, P.I. Sarafoglou b, N. Aravas b,c, G.N. Haidemenopoulos b

a Institute for Research & Technology — Thessaly, Centre for Research & Technology Hellas (CERTH), 38333 Volos, Greeceb Department of Mechanical Engineering, University of Thessaly, Pedion Areos, 38334 Volos, Greecec International Institute for Carbon Neutral Energy Research (WPI-I2CNER), Kyushu University, 744 Moto-oka, Nishi-ku, Fukuoka 819-0395, Japan

a r t i c l e i n f o a b s t r a c t

Article history:Received 2 April 2014Received in revised form 14 July 2014Accepted 26 July 2014Available online 7 August 2014

Keywords:Creep damagePigtailRefinerySteam reformer unitFE analysis

Pigtail tubes constitute a critical component of the outlet system of steam reformer unitsfor hydrogen production in refineries. Creep failure of pigtails is a common cause of down-time and potential risk to plant personnel. In this paper, finite element analyses are com-bined with experimental techniques in order to evaluate creep damage in INCOLOY 800HTpigtails of a reformer unit. A finite element model has been developed in order to investi-gate the creep behavior of pigtails under various operating conditions. A metallurgical eval-uation of creep damage has been performed on pigtail samples removed after 8.5 years ofoperation. The results indicate that the most important parameters affecting the structuralbehavior of pigtails are the operating temperature, reduction of thickness, magnitude ofcounterweights and grain size of the alloy. A comparison between the numerical calcula-tions and the metallurgical evaluation is also reported.

� 2014 Elsevier Ltd. All rights reserved.

1. Introduction

Hydrogen consumption in oil refineries is becoming increasingly important due to the sulfur reductions accomplished inthe end products and due to the addition of new hydrocracking units, which require hydrogen for their operation. Steam-hydrocarbon reforming is the most commonly used process for large-scale hydrogen production accounting for more thanð95%Þ of hydrogen consumed in oil refineries. Hydrogen is produced by a high-temperature reaction between steam andhydrocarbon gases in the presence of a catalyst, inside an array of heat-resistant, centrifugally-cast tubes (catalyst tubes).These are high Cr–Ni alloy tubes that are connected to the main transfer lines with a system of pigtail tubes, which, dueto their convoluted geometry, accommodate the relative thermal expansion between the catalyst tubes and the main trans-fer lines. A common arrangement of the outlet system is shown in Fig. 1. Failure of these outlet pigtails is a common cause ofdowntime and potential risk to plant personnel.

Prior published investigations regarding damage in pigtail tubes are limited. Roumeau [1] investigated the high temper-ature cracking of pigtails which cracked after only three (3) years of operation. The failure was attributed to creep deforma-tion which was accelerated due to the fine grain size of the material at the bends because of poor cold working of the tubes.Kodali and Ritchert [2] identified longitudinal cracks in steam reformer pigtails and attributed the failure to a combination ofcreep damage and high-temperature oxidation. Monteiro [3] discussed the failure of INCOLOY 800H pigtails by perforation

ie.uth.gr

Fig. 1. A common arrangement of the outlet system in a refinery steam reformer unit.

L.A. Spyrou et al. / Engineering Failure Analysis 45 (2014) 456–469 457

due to service at high temperature. It was shown that cracking resulted by interconnection of voids at grain boundaries andthat the curved parts were in a more advanced stage of creep compared to the straight parts. This was attributed to the smal-ler grain size of the curved parts, due to in-service recrystallization following cold bending. More recently, Thomas and Smil-lie [4] presented an overview of failure mechanisms in pigtails, the major of which are creep deformation, creep fatigue,environmental attack, and relaxation cracking. Gommans et al. [5] showed that the cracks in the bottom manifold of a steamreformer were due to strain-assisted intergranular oxidation, which was strongly influenced by grain size, and that creeplimited by grain-boundary diffusion (Coble creep) was the acting deformation mechanism for INCOLOY 800H at 800 �Cand low stresses. Xu et al. [6] investigated the failure of an INCOLOY 800HT pipe operating at 1032 �C. They determined thatthe failure was not caused by creep but from high operating stresses exceeding the yield strength of the material. A creepremaining life assessment of an INCOLOY 800H material was performed by Maharaj et al. [7] by determining the percentagecreep cavities versus the tube outer diameter. A good correlation was achieved by fitting the data points to the classic creepstrain versus time curve. A real time monitoring system of operating parameters for high temperature components, includ-ing inlet pigtails, used for the calculation of the materials remaining life fraction has been recently discussed by Daga andSamal [8].

The aim of this paper is the evaluation of creep damage in the outlet pigtails of a steam reformer unit. In the case underinvestigation the pigtails exhibited excessive deflections after seven (7) years of operation. A photo of the outlet system andthe supporting ‘‘counterweights’’ are shown in Fig. 2. The excessive deflection of certain pigtails is indicated by the dash-white-line in Fig. 2. The pigtail material was INCOLOY 800HT (UNS 08811), which is the industry standard for pigtail con-struction; it is an Fe–Ni–Cr alloy with additions of Al and Ti. The evaluation procedure includes both finite element analysisand experimental techniques. A finite element model was developed first in order to determine whether the excessivedeflections are due to extensive creep in the pigtails and to identify the locations of high creep strains where creep damageis more likely to occur. Additional analyses were carried out in order to investigate the effects of reduced pipe thickness due

Fig. 2. Deformation of the outlet system. The vertical deflection of the pigtails is shown with the up down arrow.

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to corrosion during prolonged exposure of the pigtails to elevated temperatures. When excessive deflections are observed, itis common practice in refineries to increase the counterweights; additional calculations were carried out in order to inves-tigate the effects of increased counterweights.

When the period for the planned maintenance shutdown of the plant was reached (after 8.5 years of operation) sampleswere taken from the removed pigtails for metallurgical evaluation of creep damage. The samples were taken from locationsof high potential risk to creep failure as predicted by the finite element solution.

2. Finite element modeling

Finite element calculations were carried out to determine creep-induced deflections and locations of high creep strainsin pigtails of a steam reformer unit. A typical creep behavior of a material consists of three characteristic stages: (i) theprimary creep where the strain rate is relatively high but decreases with increasing strain over a short period of time, (ii)steady-state creep where the creep strain rate remains approximately constant for the most of the material’s lifetime, and(iii) tertiary creep where the creep strain rate accelerates rapidly due to damage accumulation in the material. Primary creepstrain is usually less than one percent of the sum of the elastic, primary, and steady-state strains whereas the property that ismost commonly used to characterize creep for design purposes is the strain rate in the steady-state creep stage.

In the current study, the purpose of the numerical analysis was mainly to identify the locations of high creep strains fromwhere samples should be taken for a metallurgical assessment of creep damage and not the failure prediction and remainingcreep life of the system. Therefore, a power law creep law that represents the steady-state stage was chosen to describe thecreep behavior of the material neglecting primary and tertiary creep. A different approach including damage parameters inthe creep law should be taken if the numerical analysis was intended for the damage or lifetime prediction of the structure[9].

A single pigtail was analyzed under normal operating conditions. The geometry analyzed was based on the arrangementof the outlet system shown in Fig. 1. The pigtail tubes have an outer diameter of 42:16mm and thickness 5:8mm. The pigtailtubes are covered by insulation and an aluminum foil of thickness 70mm and 0:8mm respectively. The whole system is sup-ported by ‘‘counterweights’’ through a system of pulleys as shown in Fig. 2. The pigtails are made of INCOLOY 800HT and setto operate at a mean temperature of 835 �C under an internal pressure of 2:2MPa. The pigtail model was also analyzed con-sidering several loading and damage scenarios.

2.1. Model description

Fig. 3 shows a 3-dimensional geometrical representation of the pigtail and Fig. 4 shows the corresponding dimensions.The analysis was carried out using the ABAQUS v6.12 general purpose finite element program [10]. The pigtail was dis-

cretized using quadratic 3-node pipe elements (PIPE32 in ABAQUS). These beam-type elements, in addition to the bendingstresses, provide for a uniform radial expansion of the cross-section caused by the internal pressure and account for the cor-responding hoop stresses that develop. The nodal degrees of freedom (DOF) are the three components of displacement andthe three components of rotation, a total of 6 DOF per node. The total number of elements and DOF in the finite elementmesh are 1128 and 15,798 respectively.

Fig. 3. A 3D geometrical representation of the pigtail.

Fig. 4. Dimensions of the pigtail in mm.

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A ‘‘finite strain’’ analysis was carried out, i.e., the geometric non-linearities resulting from the geometry changes due todeformation and rotation were accounted for. The total deformation rate D is written as the sum of an elastic, thermal, andcreep part:

D ¼ De þ Dth þ Dcr: ð1Þ

The elastic part of the deformation rate De is related to stress through a linear isotropic hypoelastic equation of the form

De ¼ 1þ mE

rO

� m1þ m

_rkkd

� �; ð2Þ

where E is the elastic Young’s modulus, m is the Poisson ratio, r is the stress tensor, rO

is the Jaumann or co-rotational stressrate, and a superposed dot indicates material time differentiation. The values of E and m at two temperature values are givenin Table 1 [11].

The thermal part Dth is determined from the temperature rate _T:

Dth ¼ a _T d; ð3Þ

where T is the temperature, d is the identity second-order tensor, and a is the thermal expansion coefficient specified inTable 1 [11].

The creep deformation rate used in the calculations is a three-dimensional version of the well-known Bailey–Norton lawand has the form

Dcr ¼ _�ecrN; _�ecr ¼ Brneq; N ¼ 3

2reqs; ð4Þ

where B;nð Þ are material constants defined in Table 1 [11], req ¼ffiffiffiffiffiffiffiffiffiffiffi32 sijsij

qis the von Mises equivalent stress, s ¼ r� pd is the

stress deviator, p ¼ rkk=3 is the hydrostatic stress, _�ecr ¼ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi23 Dcr

ij Dcrij

qis the equivalent creep strain rate, and the summation

convention on repeated Latin indices has been used. The values of the B;nð Þ parameters used in the analysis correspondto a material that has not been exposed to high temperatures.

The applied loads are: (a) the internal pressure of 2:2MPa in the pigtails, (b) the distributed load due to the weight of thepigtails (mass density q ¼ 7940 kg=m3), (c) the distributed load of the insulation (qins ¼ 135 kg=m3), (d) the distributed loadof the aluminum foil (qAl ¼ 2700 kg=m3), and (e) the counterweights 1A, 2A and 3A of m1 ¼ 79:05 kg;m2 ¼ 100 kg, andm3 ¼ 194:5 kg respectively, as shown in Fig. 5. Counterweights 1A and 2A are applied at both ends of a beam supporting10 pigtails, whereas counterweight 3A is applied at both ends of a beam supporting 20 pigtails. Therefore, the counter-weights on a single pigtail are

Table 1Material data used in FE calculations for INCOLOY 800HT.

Temperature (�C) E (MPa) m a (/�C) B (/h�MPanÞ) n

835 140 0.40 18:0� 10�6 4:5� 10�20 9.0

925 130 0.41 18:5� 10�6 2:9� 10�11 4.47

F (1A)1 F2 (2A)

F3 (3A)

Fig. 5. A schematic representation of the boundary conditions and the applied loads on the pigtail.

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F1 ¼m1 g

5¼ 155 N; F2 ¼

m2 g5¼ 196N; F3 ¼

m3 g10¼ 191 N; ð5Þ

where g ¼ 9:81m=s2 is the acceleration of gravity. The counterweight loads F1; F2, and F3 are distributed over several beamelements covering a length of 150 mm each (Fig. 5).

All nodal DOF were set to zero at the two endpoints of the pigtail. A schematic representation of the applied loads and theboundary conditions is shown in Fig. 5.

The total vertical force on a pigtail due to the counterweights is directed upwards and has magnitude

F ¼ F1 þ F2 þ F3 ¼ 542 N: ð6Þ

The distributed load per unit length of the beam elements is

q ¼ R2o � R2

i

� �qþ R2

o;ins � R2i;ins

� �qins þ R2

o;Al � R2i;Al

� �qAl

h ig ¼ 96:5

Nm; ð7Þ

where Ri ¼ 15:28 mm;Ro ¼ Ri;ins ¼ 21:08mm;Ro;ins ¼ Ri;Al ¼ 91:08 mm, and Ro;Al ¼ 91:88 mm. The total length of the pigtail is‘ ¼ 5:631 m. Therefore, the total vertical load on a pigtail due to self-weight is directed downwards and has a magnitude of

W ¼ q‘ ¼ 543 N ð8Þ

It should be noted that the total counterweights are equal to the self-weight of the pigtails, i.e., F ffiW .

2.2. Simulation cases

In the following sections we study separately the effects of operating temperature, counterweights, and pigtail thicknesson the stresses that develop in the pigtails and resulting distortions. A modified material microstructure with finer grainscould result by cold working with no annealing during shaping of the elbows. The effect of grain size could be taken intoaccount in the analysis by modifying the B-parameter that enters the Bailey–Norton creep law (see Eq. (4)). In order toaccount for a material with finer grains a larger value of the B-parameter should be used; this leads to a material with accel-erated creep behavior. However, the analysis of the grain size effect was not included in the present study.

2.2.1. Analysis for operating temperatures of 835 �C and 925 �CThe mean operating temperature of the pigtails was 835 �C. In order to take into account possible temperature raise dur-

ing the operating life, additional analysis for an elevated operating temperature of 925 �C was also carried out.Figs. 6a and 7a show the deformed configurations of a pigtail for the two cases analyzed after 7 years of operation. As

expected, much larger distortions are predicted for the case of a continuous operation at the unrealistically high temperatureof 925 �C.

Figs. 6b and 7b show contour plots of the accumulated equivalent creep strain �ecr after seven (7) years of continuous oper-ation at temperatures T ¼ 835 �C and T ¼ 925 �C respectively. At T ¼ 835 �C the maximum creep strain appears at the upperright elbow of the pigtail, whereas at T ¼ 925 �C the maximum appears at the lower right elbow.

Fig. 8 shows the temporal variation of the reaction forces at the two supports for a continuous operation temperature ofT ¼ 835 �C. Fig. 8a shows the variation over the first six months of operation and Fig. 8b shows the seven-year variation.Fig. 8 shows an instantaneous elastic response of the reaction forces followed by relaxation due to creep deformation ofthe pigtails. The maximum value of the reaction forces appears at the start of operation (elastic response); the reactions thenrelax and change sign as creep takes places (Fig. 8), taking opposite values at all times, since the total counterweights areequal to the weight of the pigtails and the insulation. The reaction bending moments show a similar behavior and take theirmaximum values of about 270 (lower support) and 100 N m (upper support) at the start; then, they relax settling down toapproximately equal values of 75 N m.

Fig. 6. After 7 years of continuous operation at temperature 835 �C (a) deformed and undeformed configurations of the pigtail, and (b) �ecr contours.

Fig. 7. After 7 years of continuous operation at temperature 925 �C (a) deformed and undeformed configurations of the pigtail, and (b) �ecr contours.

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Fig. 8. Temporal variation of the vertical reaction forces at the two supports when the operating temperature is 835 �C (a) during the first six months ofoperation, and (b) during 7 years of operation.

Fig. 9. Vertical deflection at the lower right of the pigtail after half a year of continuous operation with actual and increased counterweights at temperature(a) 835 �C, and (b) 925 �C.

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2.2.2. Increased counterweightsWe examine next the effect of the magnitude of counterweights on the pigtail deflection. A new set of calculations was

carried out in which the counterweights 1A, 2A and 3A shown in Fig. 5 were increased from (79.05, 100, 194.5) kg to (80, 105,200) kg respectively. Fig. 9 shows the temporal variation of the vertical deflection of the point where lower right counter-weight F2 was applied (point 2A in Fig. 5) for the two temperatures analyzed. For both T ¼ 835 �C and T ¼ 925 �C, use ofincreased counterweights results in reduced deflection, the effect being more pronounced at the elevated temperature of925 �C. At the normal operating temperature of 835 �C, most of the deflection appears at the start of operation (instanta-neous elastic response) and the subsequent effects of creep deformation are less significant. At the extreme case of 925 �Cthe creep effect is substantial.

2.2.3. Reduced thickness analysisIn this section the effects of the pigtail wall-thickness are examined. In a new set of calculations the thickness of the pig-

tail tube was reduced from 5:8 mm to 4:8 mm in order to account for possible corrosion of the pigtails due to their exposureat high temperatures.

Contour plots of the equivalent creep strain �ecr after 7 years of operation are shown in Fig. 10. In the case of T ¼ 835 �C,where the strains are small, there is little difference between the new results and those of the original thickness of 5:8 mm;however, the strains are substantially larger in the case of the higher temperature T ¼ 925 �C.

Fig. 11 shows the temporal variation of the vertical deflection of point 2A of Fig. 5, where counterweight F2 was applied.Again, there seems to be very little effect at T ¼ 835 �C, whereas an accelerated deflection is predicted at T ¼ 925 �C, whenthe thickness is reduced to 4:8mm.

Fig. 10. Contours of �ecr after 7 years of continuous operation with lower pipe thickness 4:8 mm at temperature (a) 835 �C, and (b) 925 �C.

Fig. 11. Vertical deflection at the lower right of the pigtail after half a year of continuous operation with actual and lower pipe thickness at temperature (a)835 �C, and (b) 925 �C.

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3. Experimental study

In the case under investigation the pigtails of the steam reformer unit were removed after 8.5 years of operation, whenthe period for the planned maintenance shutdown of the plant was reached. Samples were taken from two of the removedpigtails for metallurgical evaluation of creep damage. The samples were taken from locations where the finite element solu-tions predicted maximum and minimum values for the equivalent creep strain �ecr .

3.1. Procedures

Specimen selection: Pigtail tube samples consisting of both curved and straight parts from two pigtails (P1 and P2) wereselected for examination. As shown in Fig. 12, two straight parts O and A were selected. Part O corresponds to the connectionwith the catalyst tube while part A corresponds to the connection with the main transfer line. The curved parts B and C

Fig. 12. Parts from pigtail samples selected for examination.

Table 2Chemical analysis of pigtail material (Alloy 800HT).

Ni Cr Mn Si Al Ti C

Present 31.8 22.0 1.15 0.57 0.44 0.48 0.1ASTM B407-01 30–35 19–23 1.5 max 1.0 max 0.15–0.6 0.15–0.6 0.06–0.1

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correspond to the upper and lower bends respectively. Several specimens were prepared from the pigtails for the followinglaboratory analyses:

Chemical analysis: Chemical analysis was carried out in the O2 sample, employing the method of optical emission spec-trometry analysis. The results are shown in Table 2 and are within the limits set by the ASTM B407-01 specification.

Stereo optical microscopy: Stereo-Optical microscopy was performed on the outer surface of the tubes. A Leica ‘‘Wild M3Z’’stereo-optical microscope was used at magnifications 6.5�–40�.

Metallography: Full metallographic analysis was performed on transverse (T) and longitudinal (L) cross sections of thetubes. Specimen preparation included cutting with Struers ‘‘Accutom 2’’, mounting of the specimens in resin, grinding withSiC papers 120, 320, 500, 800 and 1000 grit, and polishing with diamond paste of 3 and 1 lm diameter. Etching was per-formed with a mixture of HCl and H2O2. Examination of the metallographic specimens was carried out on an Optical Metal-lographic Microscope, Leitz ‘‘Aristomet’’ at magnifications 50�–1000�. Grain size was determined according to ASTM E112.

Microhardness measurements: Microhardness measurements were performed on selected metallographic specimens.Microhardness tester WOLLPERT 402MVD was used at a load of 300 g on Vickers scale.

Scanning Electron Microscopy: Selected specimens were examined with a SEM JEOL JSM-5310. Local chemical analysesemploying EDX were also carried out.

3.2. Metallurgical assessment of creep damage

Pigtails P1 and P2 exhibited similar creep behavior, therefore the results of the metallographic evaluation of creep dam-age are collected in Table 3 for both pigtails. Creep damage was classified in three categories.

� Isolated cavities – initial stage of creep.� Oriented cavities – intermediate stage of creep.� Microcracks – advanced stage of creep.

The curved parts C and the straight part A exhibited the advanced stage ‘‘microcracks’’. The curved part B exhibited thestage ‘‘oriented cavities’’ while the straight part O exhibited only the initial creep damage in the form of ‘‘isolated cavities’’.The grain size generally lies in the range of ASTM No 4–5. In the curved parts the grain size was not homogeneous and smal-ler than the straight parts reaching size ASTM 6. The recommended specification for alloy 800HT is ASTM 5. The difference ofgrain size between the curved and straight parts is attributed to in-service recrystallization. This recrystallization took placebecause annealing was not carried out after the cold bending of the tubes in the assembly stage of the reformer. Microhard-ness was measured through the tube thickness. A uniform microhardness profile from internal to the external surface of the

Table 3Metallographic evaluation of creep damage.

Isolated cavities Oriented cavities Microcracks

PartsO_T U

O_L U

B_T U U

B_L U U

C_T U U U

C_L U

A_T U

A_L U U U

Fig. 13. Two types of precipitates: M23C6 carbides and Ti(C,N) carbonitrides.

Fig. 14. Microstructure of transverse section of spec O1 T. Isolated cavities are indicated with arrows.

L.A. Spyrou et al. / Engineering Failure Analysis 45 (2014) 456–469 465

tubes was obtained with values in the range 145–155 HV0:3. Representative micrographs from the metallographic evaluationare shown in Figs. 13–18.

In general the microstructure contains several types of precipitates. Carbides (mostly Cr carbides, M23C6) were observedmainly at grain boundaries and triple points and also within the grains (Fig. 13). Formation of carbides at grain boundaries isthe precursor of creep damage. Ti(C,N) particles were also observed in the microstructure. Formation of titanium carbonit-rides, Ti(C,N) is common in Alloy 800HT pigtails. Ti(C,N) are usually formed during manufacture of the pigtail material andare generally etch as gold to orange precipitates within the microstructure. Fig. 14 corresponds to sample O1 T (transversesection of straight part O in pigtail P1) and exhibits isolated cavities which are mostly related to particles at grain boundaries.

Fig. 15 corresponds to sample B1 T (transverse section of curved part B in pigtail P1) and exhibits oriented cavities, whichare aligned in a straight line. Fig. 16 corresponds to sample A1 L (longitudinal section of straight part A of pigtail P1) andexhibits microcrack formation along the grain boundaries.

Fig. 15. Microstructure of transverse section of the spec B1 T. Oriented cavities are indicated with arrows as well as isolated cavities.

Fig. 16. Microstructure of longitudinal section of the spec A1 L. Microcrack is shown along grain boundary with an arrow.

Fig. 17. Isolated cavity in spec O1 T.

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The metallographic results were verified with SEM analysis. Fig. 17 shows an isolated cavity in sample O1 T whileFig. 18(a) and (b) shows microcrack formation in sample C1 T. In addition to creep, the pigtails exhibit a form of environ-mental attack at the external surface. The depth of attack is of the order of 200 lm (3–4 grains). A surface scale has formedwith thickness 25–30 lm, while the grain boundaries are also affected. This phenomenon appears in all pigtails examined. Acharacteristic micrograph is shown in Fig. 19 for specimen B1 L. In addition Needle-like AlN precipitates form within thisregion and are depicted in Fig. 20 for specimen O1 T.

Fig. 18. (a), (b), Microcrack formation in sample C1 T.

Fig. 19. Intergranular oxidation and decarburization of external surface of spec B1 L.

Fig. 20. Needle-like AlN precipitates in spec O1 T.

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SEM/EDX was conducted in the region close to the external surface to identify the origin of environmental damage. Fig. 21is an EDX spectrum from the matrix to be used as reference. In Fig. 22 the spectrum from the scale, indicates that it is Croxide. Fig. 23 is a spectrum from the grain boundaries indicating segregation of alloying elements Ti, Al, Si and Cr andsubsequent intergranular oxidation.

Finally Fig. 24 is a spectrum from a subscale particle rich in Cr, indicating Cr oxide with segregation of Ti and Al. The roomtemperature tensile mechanical properties were measured in the straight part O of the pigtails. The results are shown in

Fig. 21. EDX spectrum from the matrix to be used as reference.

Fig. 22. Spectrum from the scale, indicating that it is Cr oxide.

Fig. 23. Spectrum from a grain boundary.

Fig. 24. Spectrum from a subsurface particle.

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Table 4Tensile mechanical properties of the pigtail material (Alloy 800HT).

Yield strength (MPa) UTS (MPa) Elongation, 2 in, 5D ð%Þ

Present 415 593 25ASTM B407-01 170 min 450 min 30

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Table 4 and are compared with the specified values for the unexposed material [12]. Compared with the ASTM specificationthe strength values are within limits and there is a slight decrease of the elongation, which is considered normal after8.5 years exposure at 835 �C due to aging phenomena.

4. Discussion and conclusions

Under normal operating conditions (835 �C) high creep strains are predicted to appear at parts ‘‘A’’ and ‘‘B’’ (Fig. 10a).When the pigtails are exposed to higher temperatures (925 �C), Fig. 10b shows that larger deflections develop and high creepstrains appear at parts ‘‘A’’ and ‘‘C’’. The metallurgical evaluation presented above has shown that higher creep damage in theform of microcracks has appeared in parts ‘‘A’’ and ‘‘C’’; this is in agreement with the FEM calculations for the case of elevatedtemperature exposure. The observed higher deflections shown in Fig. 2 are due to exposure of the pigtails at a temperaturehigher than the normal operating temperature of 835 �C (see Fig. 10a and b).

The most important parameters affecting the structural behavior of pigtails at high temperatures are:

� Operating temperature.� Reduced thickness due to corrosion.� Magnitude of counterweights.� Grain size of alloy.

In particular higher operating temperature and reduced thickness due to corrosion can have significant effects on theobserved deflection, creep strains, and creep damage. The effect of counterweights in decreasing the observed deflectionis important at high operating temperatures. Grain size of the alloy is also an important parameter, since higher creep resis-tance is associated with a larger grain size. The recommended grain size for INCOLOY 800HT is ASTM No 5. However parts‘‘B’’ and ‘‘C’’ exhibited a smaller grain size of No 6 and actually creep damage in these parts was higher than the rest of thepigtail.

Acknowledgments

The authors are wish to acknowledge the assistance of Mr. P. Dimitriadis and Mr. I. Altanis of Motor Oil Hellas forproviding pigtail specimens. The assistance of Dr. H. Kamoutsi in the preparation of the figures is also appreciated.

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