Post on 24-Apr-2020
2011 [83] AMMONIA TECHNICAL MANUAL
Mechanical Integrity of Primary
Reformer Hot Outlet Headers
This paper discusses mechanical integrity issues for primary reformer hot outlet headers. These are
typically welded assemblies of wrought Alloy 800H or 800HT or cast 20Cr-32Ni-Nb alloy. The typical
design and operating conditions for these headers are presented. Problems and failures that have been
reported at past Ammonia Plant Safety Symposiums are reviewed and summarized. With this
background information, recent failures of hot outlet headers are described along with the identified
causes of these failures. The most common failure mechanism is high-temperature creep but the
location and causes of creep damage vary. Creep failures have occurred both at weld joints and within
base metal. The sources of these problems are discussed and suggestions for mitigation are presented.
Based on the findings of the failure and material properties investigations, stress analysis and material
damage modeling is used to estimate remaining life.
Carl E. Jaske
Det Norske Veritas (USA), Inc., Dublin, OH USA
Greg T. Quickel
Det Norske Veritas (USA), Inc., Dublin, OH USA
Introduction
he purpose of this paper is to review and
discuss the primary mechanical integrity
issues that are associated with the hot out-
let headers of primary reformers. A number of pa-
pers presented at past Ammonia Plant Safety
Symposiums have addressed various issues with
these hot outlet headers. To provide background
for recent issues with similar headers, the past
work is reviewed and the various types of prob-
lems that have been encountered in the past are
summarized. Relevant issues on similar headers
in hydrogen and methanol reformers are also dis-
cussed. In all cases, the hot outlet headers operate
in the high-temperature creep range for their ma-
terial of construction, which is either a wrought
Alloy 800H or 800HT or a cast 20Cr-32Ni-Nb al-
loy. Because of the operating temperature range,
failures typically occur by high-temperature creep
or creep-fatigue, with the fatigue damage induced
by thermal cycling. Thus, creep strength is an im-
portant parameter addressed in the review discus-
sion.
Following the background review, examples of
recent problems with hot outlet headers are pre-
T
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sented and discussed. The causes of these prob-
lems and their relation to past problems are eva-
luated. The purpose of this evaluation is to help
operators of primary reformers minimize the po-
tential of similar problems in future operation of
their units and to provide guidance for estimating
the safe and useful life of headers that are current-
ly in service. Such guidance will help avoid un-
planned shutdowns and schedule repairs or re-
placements in a timely, cost-effective manner.
Past Hot Outlet Header Issues
Steam-methane reformers have either hot or cold
outlet headers to collect the process gas from the
outlet of the catalyst tubes. Cold headers are made
of carbon steel and insulated so that they operate
at temperatures well below the creep range, while
hot headers are not insulated and operate at high
temperatures close to the reformer outlet gas tem-
perature. The focus of this paper is on hot outlet
headers that operate in the high-temperature creep
range of the materials from which they are con-
structed. Kawai, et al. [1] point out that there are
two basic types of hot header system designs: (1) a
hot collector where the catalyst tubes are attached
directly to the manifold by means of fittings and
the hot gases exit from the manifold via riser
tubes and (2) a hot subheader with a cold transfer
line where the catalyst tubes are attached to the
subheader via outlet pigtails. In either case, the
hot header or subheader is cylindrical vessel with
numerous welded attachments at holes that allow
the hot process gas to flow into the header and
central tee section where the hot process gas flows
out of the header.
Because the headers operate at temperatures near
or greater than 800 °C, they are typically made
from either wrought Alloy 800Hor 800HT or from
cast 20Cr-32Ni-Nb alloy. The fittings and tees are
also made using these same alloys. Welding of the
headers, tees, and fittings is done with a high-
nickel material or a “matching” material. Com-
monly used welding electrodes include Inco-Weld
A, Inconel 182, Inconel 112, and Inconel 117.
Commonly used filler metals include Inconel 82,
Alloy 625, and Alloy 617. The common “match-
ing” filler metals are Thermanit 21/33, Thermanit
21/33 So, and UTP A 2133 Mn. The final selec-
tion of welding material depends on the design
and the welding technique that is used.
The primary load on outlet headers is internal
pressure but bending loads are often induced by
movement of attachments and/or restraints at sup-
ports. These system bending loads are related to
thermal expansion and complex interactions with-
in the manifold system, making stress analysis to
compute their magnitude difficult and adding sig-
nificant uncertainty to the results of the calcula-
tions. Furthermore, the bending loads are often
displacement controlled so that stresses induced in
the high-temperature creep regime can relax and
redistribute with time at operating temperature. In
contrast, internal pressure is usually maintained at
a relatively constant level and is a sustained load
so that stresses induced by internal pressure are
relatively straight forward to compute.
This section of the paper reviews hot outlet header
issues that have been discussed in previous publi-
cations. Issues with primary ammonia reformers
as well as issues with similar components in hy-
drogen and methanol reformers are addressed. In
comparing this information, keep in mind that
these different types of steam-methane reformers
operate in different ranges of temperature and
pressure as illustrated in Figure 1. Ammonia re-
formers operate at higher pressures and lower
temperatures than hydrogen or methanol refor-
mers. This difference in operating conditions af-
fects the hot outlet header design. Thus, the outlet
headers in ammonia reformers usually operate at
higher stresses and lower temperatures than those
in hydrogen and methanol reformers.
Performance of Header Materials
In 1963, Francis and Glass [2] reported the failure
of outlet piping and described high-temperature
stress-rupture failures in Type 304 stainless steel
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piping. Avery and Valentine [3] reviewed the de-
sired characteristics of materials for outlet mani-
fold systems, while Harnby [4] pointed out that an
outlet manifold system made of Alloy 800 had
better creep resistance than one made of the HU
alloy. Later, Blackburn [5] pointed out that mani-
folds made of Alloy 800 were less prone to crack-
ing than those made of traditional cast alloys, such
as HU, HK, and HT, because of its superior duc-
tility. He proceeded to note that the cast 20Cr-
32Ni-Nb alloy was an improvement to Alloy 800
for hot outlet header applications. These papers
provide historical background on alloys for use in
hot outlet headers.
0
2
3
4
700 800 900 1000
NH3
H2
CH3OH
Limit of Tube
Metal Use
1
Gas Temperature, °C
Internal Pressure, MPa
0
2
3
4
700 800 900 1000
NH3
H2
CH3OH
Limit of Tube
Metal Use
1
0
2
3
4
700 800 900 1000
NH3
H2
CH3OH
Limit of Tube
Metal Use
1
Gas Temperature, °C
Internal Pressure, MPa
Figure 1. Typical operating temperature and pres-
sure ranges for ammonia (NH3), hydrogen (H2), and
methanol (CH3OH) reformers, from Kawai, et al. [1].
Shibisaki, et al. [6] found that Si, Nb, and C addi-
tions had a negative effect on the ductility of
20Cr-32Ni-Nb cast alloys while Mn had a positive
effect on their ductility. Parks and Schillmoller [7]
indicated that Alloy 800H has a low tendency for
embrittlement but noted that the 20Cr-32Ni-Nb
cast alloy has higher creep-rupture strength than
Alloy 800H. Thomas, et al [8] pointed out a se-
rious embrittlement issue with the 20Cr-32Ni-1Nb
cast material and suggested that niobium-free
grades should be developed in order to avoid such
in-service problems. Hoffman [9] evaluated the
aging of 20Cr-32Ni-Nb castings after 21 months
of service at 852 °C and found both M23C6 carbide
and inter-metallic Nb-Ni silicide precipitates
present in the aged material. The silicides were
mostly G phase plus some η’ phase. They are ex-
pected to decrease creep resistance and promote
liquation cracking during welding.
Hoffman and Gapinski [10] evaluated samples of
a 20Cr-32Ni-Nb cast outlet manifold that had de-
veloped fatigue cracks after it had been in service
for approximately 192 months at an outlet gas
temperature of 871 °C. In addition to carbide and
silicide precipitation, they observed carbide coa-
lescence. The observed aging significantly de-
graded the tensile strength, Charpy V-notch im-
pact resistance, stress-rupture strength and
weldability of the material. The carbides and sili-
cides were solutionized by annealing at 1093 °C.
However, complete solutionizing of secondary
precipitates and spheroidization of the primary
carbides required a minimum annealing tempera-
ture of 1149 °C. Better stress-rupture properties
were achieved by solution annealing at 1204 °C
than by solution annealing at 1093 or 1149 °C.
The cracked region of the manifold was success-
fully repaired after in-situ solution annealing at
1149 °C for three hours and operated successfully
for four more years before the furnace was rebuilt.
The authors recommended keeping Si content as
low as possible and maintaining a stoichiometric
Nb/C ratio of 7.7.
Welding and Repairs
Roach and VanEcho [11, 12] found that weld-
ments generally have lower stress-rupture strength
than comparable base metals. Inconel 182 weld-
ments had the lowest stress-rupture strength when
compared with Inconel 82, Inconel 112, and Inco-
Weld A weldments. Several reformer operators
[13-15] have experienced accelerated creep dam-
age, embrittlement, and failures with Inconel 182
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weld metal. Failures have been attributed to a loss
of rupture strength caused by embrittlement. Inco-
nel 182 is now known to exhibit a loss of rupture
strength and ductility at elevated temperatures.
Commonly used materials for joining Alloy 800
are the Inco-Weld A electrode and Inconel 82 fil-
ler metal. Avery and Valentine [5] specifically
recommended filler metal Inconel 82 for Alloy
800 welds. Roach and VanEcho [11, 12] indicated
that Inconel 82 weldments have higher stress-
rupture strength than Inco-Weld A and Inconel
182 weldments, but lower stress-rupture strength
than Inconel 112 weldments. McMillan [13]
found success in using Inco-Weld A material in
place of failing Inconel 182 filler metal.
Shibisaki, et al. [16] recommended the use of a
high-nickel alloy metal (Inconel 82) for welding
20Cr-32Ni-Nb metals. Kobrin [15] obtained supe-
rior results using Inconel 82, Alloy 625 filler met-
al, and Inconel 112 in place of Inconel 182. Or-
bons [17] suggested using Inconel 112 in place of
Inco-Weld A for applications at temperatures
above 816 °C because of its superior resistance to
creep damage. Roach and VanEcho[11, 12] also
found that Inconel 112 had superior creep resis-
tance to those of Inco-Weld A, Inconel 182, and
Inconel 82 weldments. van Wortel [18] noted that
filler metals 82 and 617 were found to be suscept-
ible to relaxation cracking. For applications that
require the highest strength and corrosion resis-
tance, Inconel 117 electrode and Alloy 617 filler
metal are recommended [19] for use at tempera-
tures above 788 °C.
Al-Jubeihi [20] reported making successful repair
welds by grinding out cracks, depositing Inco-
Weld A stringer beads, and grinding out welds
and base metal defects one bead at a time while
using dye penetrant tests and ultrasonic examina-
tions to ensure quality repairs. Another successful
repair method [21] involves grinding into through-
wall cracks to remove 60% of the cross-section
and grinding back the surface area by 5 mm to
prevent re-cracking associated with embrittlement.
The beveled area is then covered using Inconel
182 electrodes, and the previously ground grooves
are filled up by welding following a stringer bead
technique. The shortcoming of this repair method
is that Inconel 182 welds have poor high-
temperature creep strength.
Pande and Swenson [22] successfully completed a
preliminary repair by welding a reinforcement
cone onto sound material in the tees to extend op-
erating life. Blackburn [5] developed a method to
improve weld efficiency using parent metal as fil-
ler after completing the root weld with Inconel 82
wire.
Shi and Lippold [23] found that repair practices
that lead to failures often involve elevated temper-
atures throughout welding repairs leading to em-
brittlement, which in turn gives rise to a ductility
dip. When the elevated temperature restraint ex-
ceeds the ductility minimum, failure occurs.
Another common problem is the mechanism of
hydrogen induced crack formation which leads to
weld failures. Pattabathula, et al. [24] noted this
can be mitigated by increasing the temperature at
the weld above dew point, by adding outside sur-
face insulation, and by installing windshields on
sides of the furnace.
Another problem with welding operations, pointed
out by Shibisaki, et al. [25], involves differences
between weld and base metal. The difference in
thermal expansion coefficients can lead to elastic-
plastic strain of the weld material. Orbons [17]
suggested high-temperature preheating around the
weld to prevent cracking of the base metal caused
by weld shrinkage. Lobley, et al. [26] found that it
was necessary to bore machine and solution an-
neal a 20Cr-32Ni-Nb cast header prior to repair
welding because of its severe embrittlement in the
aged condition, especially near the inner surface.
Anderson and Bruno [27] reviewed weldability is-
sues that they encountered during the welding of a
new cone piece to an existing tee (aged material)
in a hydrogen reformer outlet manifold. Both
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pieces were 20Cr-32Ni-Nb castings from the same
foundry. Metallurgical analysis of samples of the
previously installed, aged cone piece on which an
Electrode 117 weld bead had been deposited re-
vealed both heat-affected zone (HAZ) liquation
cracking (LC) and reheat cracking. As shown in
Figure 2, the liquation cracking occurred near the
fusion zone (FZ) and was believed to be caused by
low melting point silicate precipitates that had
formed during service. The reheat cracking was
believed to have been caused by a combination of
mismatched mechanical properties and high inclu-
sion content. A full solution annealing of the aged
material was required to weld it successfully.
Figure 2. Liquation cracking in heat-affected zone
(HAZ) of aged 20Cr-32Ni-Nb casting, from Anderson
and Bruno [27].
Penso and Mead [28] discussed cracking that they
found on existing 20Cr-32Ni-Nb statically cast tee
pieces when a new centrifugally cast 20Cr-32Ni-
Nb reducer cones were welded to them. The loca-
tion of this cracking is shown in Figure 3. It was
found to be associated with intermetallic silicate
precipitates in the existing (aged) material. The
tees were cut out and solution annealed at
1150 °C, followed by machining out of all of the
cracks. Welding was done using a GTAW process
with Alloy 617 filler wire, a 50 °C preheat, and a
maximum interpass temperature of 150 °C. A but-
tering layer was first applied using low amperage.
The joints were then tack welded, and a root pass
was applied. DP inspection was performed fol-
lowing every second fill pass and after blend
grinding the cap pass.
Figure 3. Location of cracking in aged 20Cr-32Ni-
Nb cast tee piece after welding a new cone piece to it,
from Penso and Mead [28].
Jack [29] reported cracking in a 21/33 weld be-
tween a tee and a subheader in a cast 20Cr-32Ni-
Nb outlet manifold after 14 years of service at ap-
proximately 793 °C. This cracking is shown in
Figure 4. Attempts to grind out and repair the
cracking were not successful, so tapered sleeves of
Alloy 800HT were designed and welded over the
cracked weld and similar welds in the outlet mani-
fold.
Figure 4. Cracking at subheader to tee weld; both
are 20Cr-32Ni-Nb castings, from Jack [29].
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Dejaeger, et al. [30] discussed cracks in an Alloy
800H outlet header. In this case, the header was
manufactured using a solution annealed Alloy
800H plate material that was cold formed into a
cylinder and then longitudinally welded with Ni-
base filler. No heat treatment was applied after the
manufacturing process. After two years, a leak
was detected. Examination revealed the cause to
be high residual stresses from fabrication (cold-
working/welding) or resulting from cyclic defor-
mation and accompanying relaxation processes.
Stress relaxation cracking was controlled in the
replacement header by a stabilization heat treat-
ment at 950 °C performed before welding. After
welding, a second heat treatment at 850 °C was
performed to decrease weld residual stresses. This
method successfully controlled the stress relaxa-
tion cracking in the Alloy 800H material. Howev-
er, this stabilization heat treatment temperature
was less than the recommended value [18] of
982 °C, and the post-weld heat treatment (PWHT)
temperature was less than the recommended val-
ue [31] of 900 °C.
Brongers and Jaske [32] evaluated an Alloy 800H
outlet manifold that had welds made using weld-
ing Electrode 112. It was removed from service
after normal operation at approximately 870 °C
for 64,000 hours because nondestructive testing
indicated cracks in the circumferential butt welds.
Creep-rupture tests were performed at stress levels
comparable to the service stress and temperatures
above the design temperature were selected to ac-
celerate the testing. The results showed that the
base metal did not suffer significant creep damage
from the prior service exposure and that the weld
metal had a very low resistance to creep damage.
Metallographic analysis of a weld sample taken
from a comparable in-service manifold supported
this conclusion, and data for new material re-
ported in the literature support the finding of rela-
tively low high-temperature creep strength for
welds made using welding Electrode 112. The au-
thors recommended using welding Electrode 117
for repairs.
Recent Hot Outlet Header Issues
To highlight some recent issues with hot outlet
headers, three case histories are reviewed. The
first two cases cover outlet headers from methanol
reformers, while the third case is for an outlet
header from a hydrogen reformer.
Case 1. Metallurgical Analysis of a Header
to Bull Tee Weld from a Methanol
Reformer
A metallurgical analysis was performed on a bull
tee from a methanol reformer outlet header. The
tee was connected to two header tubes and an in-
creaser cone. The tee was removed when leaks
were discovered at the girth welds after approx-
imately 50,000 hours of service.
The header, tee, and increaser cone were made of
a cast alloy 20Cr-32Ni-Nb alloy. The welds used
to join these components were made using an In-
conel 82 root pass, Inconel 82 or Inconel 182 rein-
forcement, and Inconel 182 for the remainder of
the weld. Thus, the majority of the weld filler
metal in the joints was Inconel 182.
Approach. The following steps were performed
for this analysis. The bull tee was visually in-
spected and photographed in the as-received con-
dition. Liquid dye penetrant testing was performed
on the three large diameter girth welds of the tee.
Cross-sections were removed from cracks in the
welds, mounted, polished, and etched. Light pho-
tomicrographs were taken to document the mor-
phology of cracking and microstructures. Samples
for chemical analysis were removed from the re-
ducer cone, header tube, and tee and weld metal
between the increaser cone and tee to determine
the chemical composition.
Results and Discussion. Figure 5 is a photograph
of the bull tee. Pigtails extended vertically from
fittings attached to the tee and indicate the top of
the tee. Girth welds attached the tee the increaser
cone and tee to two header tubes. The outside di-
ameter (OD) of both header tubes was 345 mm.
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The wall thicknesses of the of the header tubes
were measured at the 12:00, 3:00, 6:00, and 9:00
o’clock orientations. The average wall thickness
of the header tube was 21.4 mm.
Figure 5. Photograph of bull tee following liquid dye
penetrant testing of the girth welds.
Figure 5 also shows the tee after performing liquid
dye penetrant testing of the three welds. The red
dye shows the locations of indications. Four loca-
tions that contain indications were identified. The
indications were located at: 1) both header to tee
girth welds, 2) a pigtail to header tube girth weld,
and 3) the tee to increaser cone girth weld.
Figure 6 is a photograph of the crack-like indica-
tion located at the toe of a header to tee girth
weld. The indication is approximately 122 mm
long and present from the 1:15 to 2:30 orientation.
The liquid dye penetrant testing clearly showed
the locations of the crack indications at the girth
welds.
Figure 7 is a photomicrograph of the mounted
cross-section (Mount W3) that was removed a
header to tee girth weld (GW 3). The dashed line
in Figure 6 indicates the location where the cross-
section was removed and the arrow indicates the
polishing direction. Cracks are present in the HAZ
of the tee, HAZ of the header tube, and in the
weld metal. Although difficult to see, a columnar
microstructure is also present in the header. Fig-
ure 8 is a photomicrograph of Mount W3 showing
the HAZ of the tee and the weld metal. Creep
cracks are present at the grain boundaries in the
HAZ of the tee. Creep voids and aligned creep vo-
ids are present in the weld metal. Figure 9 is a
photomicrograph of Mount W3 showing the HAZ
of the header tube and weld metal. Relatively
wide, and also narrow, creep cracks are present in
the HAZ of the header tube. This photomicro-
graph was taken just below the outside surface
where the crack-like indication shown in Figure 6
was located.
Figure 6. Photograph of a crack located at girth
weld between a header and the tee; area indicated in
Figure 5.
Figure 7. Stereo light photomicrograph of a
mounted cross-section removed from the girth weld
shown in Figure 6 (Glyceriga Etchant).
Figure 6
Flow
Flow
Flow
Pigtail
Reducer
cone
Header Header
Flow
Crack Header
Polishing
direction
for cross-
section
Header
tube side Tee side
Figure 8
Figure 9
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Figure 8. Light photomicrograph of the mounted
cross-section showing the HAZ and weld metal near
the tee side; mirror image of area indicated in Figure
7 (Glyceriga Etchant).
Figure 9. Light photomicrograph of the mounted
cross-section showing the HAZ and weld metal near
the header tube side; mirror image of area indicated
in Figure 7.
Figure 10 is a photomicrograph of the header base
metal microstructure. The observed microstructure
is consistent with that expected for the 20Cr-
32Ni-Nb alloy. The figure shows primary carbides
located at the grain boundaries and smaller sec-
ondary carbides dispersed in the grains. The finer,
secondary carbides are indicative of high tempera-
ture exposure. Although it is difficult to be certain
since the appearance of the pre-service micro-
structure is not known, it appears that the network
of primary carbides was reduced and that they are
slightly blocky, which is also indicative of high-
temperature exposure.
The most severe creep damage was on the header
tube side rather than the tee side. The intergranu-
lar crack path is consistent with a high-
temperature creep mechanism. The finer, second-
ary carbides and blocky primary carbides are in-
dicative of high-temperature exposure.
Figure 10. Light photomicrograph of the header
base metal microstructure (Glyceriga Etchant).
The results of the chemical composition analysis
conducted on a sample removed from the base
metal of the header are shown in Table 1. The re-
sults of the analyses are consistent with the no-
minal chemical composition specifications for the
alloy at the time of its manufacture.
The results of the chemical composition analysis
conducted on a sample removed from the weld
metal between the increaser cone and tee are
shown in Table 2. Except for the nickel plus co-
balt (Ni + Co), manganese (Mn), iron (Fe), and
chromium (Cr) content, the results of the analysis
are consistent with specified values for Inconel 82
and 182. The Mn and Cr content are consistent
with the Inconel 82 specification and not with In-
conel 182 specification. The Ni + Co and Fe con-
tent are consistent with the Inconel 182 specifica-
tion and not the Inconel 82 specification.
HAZ Weld metal
HAZ
Weld metal
Primary
carbide
Secondary
carbide
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Table 1. Results of chemical analysis of samples re-
moved from the tee, header tube, and increaser cone
compared with the nominal chemical composition for
the 20Cr-32Ni-Nb alloy.
Element Tee
Wt. %
Header
Wt. %
Increaser
Cone
Wt. %
Nominal1,
Wt. %
Carbon 0.12 0.15 0.14 0.10 – 0.15
Manganese 1.05 1.01 0.77 1.50 (max)
Silicon 0.63 0.50 0.65 1.50 (max)
Chromium 20.62 19.91 20.01 19 – 25
Nickel 33.03 32.61 33.02 31 – 33
Niobium 1.17 1.08 1.05 0.50 – 1.50
Phosphorus 0.019 0.016 0.017 0.050
(max)
Iron Balance Balance Balance Balance
Table 2. Results of chemical analysis of a sample
removed from the weld between the tee and increaser
cone compared with the nominal chemical composi-
tion for Inconel 82 and 182.
Element
Weld
Wt. %
Inconel Filler
Metal 82 2
Wt. %
Inconel Welding
Electrode 182 2
Wt. %
Nickel + Cobalt 65.9 67.0 (min) 59.0 (min)
Carbon 0.037 0.10 (max) 0.10 (max)
Manganese 2.74 2.5 – 3.5 5.0 – 9.5
Iron 8.52 3.0 max 10.0 max
Sulfur 0.006 0.015 (max) 0.015 (max)
Silicon 0.17 0.50 (max) 1.0 (max)
Chromium 20.05 18.0 – 22.0 13.0 – 17.0
Niobium + Tantalum 2.42 2.0 – 3.0 1.0 – 2.5
Phosphorus 0.003 0.030 (max) 0.030 (max)
The fractional increase in Ni content would not be
expected to have a substantial effect on creep re-
sistance. The differences in the Mn, Fe, and Cr
content are likely related to dilution or enrichment
that can occur with the various passes of the Inco-
nel 82 and 182 welding materials.
Summary. A summary of the key observations
follows:
• Creep voids were present in the heat af-
fected zone (HAZ) and weld metal of the
header to tee girth weld but not in the base
metal.
• Creep cracks were present in the HAZ and
weld metal at one girth weld.
• Intergranular creep cracks penetrated the
outer surface in the HAZ and weld metal,
near the toe of the girth welds.
• The fine, dispersed secondary carbides are
indicative of high-temperature exposure.
• The microstructures of the base metals
were consistent with those expected for
20Cr-32Ni-Nb cast alloy.
• Except for Ni compositions for the tee and
reducer cone of 0.03 and 0.02 weight %
greater than specified, the base metal
composition was consistent with specified
values for the alloy.
• The composition of the girth weld was
consistent with that of Inconel 82 or 182,
and in some cases, with both specifica-
tions.
The cracking is consistent with a high temperature
creep mechanism associated with the low creep
strength of the filler material. This type of failure
can be mitigated by using a filler metal with im-
proved creep strength in the weld joint such as a
“matching” filler or and Inconel 617 filler.
Case 2. Metallurgical Analysis of Outlet
Header from Methanol Reformer
A metallurgical analysis was performed on two
sections from a methanol outlet header that leaked
in-service. One of the sections contained the leak
as shown in Figure 11 and the other (non-failed)
was located away from the leak in the same head-
er. The header samples were removed from ser-
vice after approximately the minimum design life
of 100,000 hours. The header tube sections were
centrifugally castings made of a 20Cr-32Ni-Nb al-
loy.
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Approach.Approach.Approach.Approach. The following steps were performed
for the analysis. The header tube sections were vi-
sually inspected and photographed. A ring section
was removed from the through-wall portion of the
header tube section; the other header section was
already in the form of a ring. The two ring sec-
tions were polished and light photomicrographs
were taken to document the observed creep dam-
age. The ring sections were then etched in Adlers
reagent to reveal their macrostructures and photo-
graphs were taken. A small portion of each ring
was then removed, mounted, polished, and etched
with glyceregia for detailed examination. Photo-
micrographs were taken to document the micro-
structure and maximum service temperatures were
estimated.
Using elastic finite-element stress analysis, the
maximum tensile hoop stress in the region of the
header failure was calculated to be 10.2 MPa.
Failure stresses for average stress-rupture lives of
100,000 hours (11.4 years) were computed for
header operation at various temperatures using the
manufacture’s Larson-Miller parameter relation-
ship for the 20Cr-32Ni-Nb alloy. A temperature of
963 °C was calculated to cause stress-rupture fail-
ure in 100,000 hours. This was 73 °C above the
reported maximum header operating temperature
of 890 °C. Based on this result, it is likely that the
material was somewhat weaker than indicated by
the manufacture’s data and that it was overheated
to some degree at the failure location.
Results and Discussion. Figure 11 is a photo-
graph of the as-received header tube section
(Header Section 1) that failed. The figure shows
the through-wall portion of the failure. Secondary
cracks are also visible in the vicinity. The cracks
were present from the right side of the tube sec-
tion up to approximately 174 mm from the right
side. Also visible is a bulge in the tube where the
cracks are present. The maximum diameters of the
bulged and non-bulged regions of Section 1 were
174.0 mm and 171.0 mm, respectively.
Figure 11. Photograph of the header tube section
(Header Section 1), as received.
Figure 12 is a photograph of the ring from Section
1, following polishing and etching. The location
of ring Section 1 is indicated in Figure 11. A co-
lumnar dendritic macrostructure was visible on
both ring sections.
Figure 12. Photograph of Ring Section 1 from
Header Section 1; dashed lines in Figure 11 indicate
the location of the ring (Adlers reagent).
Measured dimensions of the tube sections are
listed in Table 3 along with reported nominal di-
mensions of the tube when it was new. The no-
Through-wall
Creep
fissures
Through-wall
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minal inside diameter (ID) of the new header tube
was not reported, so it was calculated based on the
reported minimum outside diameter (OD) of a
new header tube minus twice the minimum sound
wall thickness. The strain values of the header
tube sections were based on calculated minimum
ID values. There was more strain at the bulge in
the header tube than in Header Section 2, which
was to be expected. Since baseline values of orig-
inal dimensions are not available for each tube,
the strains are only approximate because of inhe-
rent variations in the dimensions of new tubes.
Table 3. Measured dimensions of header tube sam-
ples compared to minimum values reported by client.
Description Outside
Diameter
(OD)
Wall Thick-
ness (WT)
Inside
Diameter
(ID)
Strain*
mm in mm in %
Minimum
new tube 168.3 6.625 15 0.591 —
Bulge on
Header Sec-
tion 1
174.0 6.850 15.0 0.592 4.10
Header Sec-
tion 2 172.0 6.77 15.4 0.607 2.08
* Strain based on ID minimum and measured values. ID
calculated based on OD and WT minimums and measured
values.
Figure 13 is a photomicrograph of Ring Section 1
adjacent to the OD surface. Relatively wide and
narrow creep cracks that appeared to follow the
dendritic columnar structure were present for the
full wall thickness. A low density of creep voids
was found in the cross-section. The density of vo-
ids was not high adjacent to the ID or OD surface.
The absence of creep voids near the cracks is evi-
dence of short-term overheating. Note that evi-
dence of austenite grains and twinning was appar-
ent at the ID where the through-wall cracks were
present. This suggests that the ID was cold
worked during fabrication. It is not known if this
contributed to the failure.
Figure 14 is a photomicrograph of Ring Section 2
adjacent to the OD surface. Aligned creep voids
(not linked) were present adjacent to the OD and
ID and appeared to be more prevalent adjacent to
the OD. The residual life of this tube section is
approximately 70% based on metallurgical analy-
sis and guidelines of past studies [33] of reformer
tubes. For 11 years of prior operation, this result
implies approximately 26 years remaining life, as-
suming that future operating conditions are the
same as those of the past.
Figure 13. Light photomicrograph of Ring Section 1
adjacent to the OD surface, near the through-wall lo-
cation.
Figure 14. Light photomicrograph of Ring Section 2
adjacent to the OD surface.
Figures 15 and 16 are photomicrographs of the
microstructures of Ring Section 1, near to and
away from a crack, adjacent to the OD. The pho-
tomicrographs were taken at the same magnifica-
tion. The network (cored structure) of primary
Aligned
creep voids
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2011 [94] AMMONIA TECHNICAL MANUAL
carbides does not appear to be intact near the
crack. A cored, lacelike structure was present
away from the crack, indicating that lower tem-
peratures were present away from the crack. Also
evident near the crack was the presence of sec-
ondary carbides and nitrides. The nitrides were not
evident away from the crack, which is to be sus-
pected. Nitrides are known to form adjacent to
fracture surfaces. Figure 17 is a photomicrograph
of the microstructure from Ring Section 2. The
cored, lacelike structure and secondary carbides
are present as was shown away from the crack in
Ring Section 1.
Figure 15. Light photomicrograph of the microstruc-
ture of Ring Section 1 near a crack and OD surface
(glyceregia etchant).
Based on the microstructures observed in the
header samples and the reference material, their
maximum service temperatures were estimated to
be:
• Ring Section 1: 870°C
• Ring Section 2: 840°C
The reference material had four times as much
carbon as the header samples, which made it diffi-
cult to estimate these values. Thus, there is uncer-
tainty regarding them.
The amount of niobium (1.91 wt% compared with
the specified range of 0.90 to 1.35 wt%) in the
header section was greater than the maximum val-
ue of the specified range. The amounts of the oth-
er elements met the manufacturer’s specification
for 20Cr-32Ni-Nb. Higher than desired niobium
content may have reduced the creep strength of
the 20Cr-32Ni-Nb alloy. This possibility was
pointed out in the previous review of published
investigations.
Figure 16. Light photomicrograph of the microstruc-
ture of Ring Section 1 away from a crack and near
the OD surface (glyceregia etchant).
Figure 17. Light photomicrograph of the microstruc-
ture of Ring Section 2 (glyceregia etchant).
Summary. A summary of the key observations
follows:
• The metallographic evidence (wide creep
cracks, low density of voids) suggests that
Header Section 1 failed by short-term over-
heating.
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• Ring Section 2 showed metallographic evi-
dence of creep damage. The remaining life of
this section was estimated to be 70%, while
the maximum service temperature was esti-
mated to be 840°C.
• Using finite-element stress analysis, the max-
imum tensile hoop stress in the region of the
header failure was calculated to be 10.2 MPa.
For this stress, a temperature of 963 °C was
calculated to cause stress-rupture failure in
100,000 hours, indicating that the material
was either overheated, weaker than expected,
or a combination of both.
Case 3. Metallurgical Analysis of Outlet
Header a Hydrogen Reformer
During an inspection, a significant increase in di-
ameter was noted in one of two headers in a hy-
drogen reformer. The headers were made of the
cast 20Cr-32Ni-Nb alloy. The portion of the re-
former that fed gas to this header was known to
have been operating at higher than desired tem-
peratures. Further inspections revealed the pres-
ence of cracks in the header, especially in the
around the holes where fittings for the outlet pig-
tails were attached. As a result, the operator de-
cided to replace the header and cut out two ring
sections for metallurgical failure analysis. Ring
Section 1 was from the central portion of the
header tube where the increase in diameter was
the largest and the most material damage was ex-
pected to be found. Ring Section 2 was from near
the end of the header tube where the increase in
diameter was the smallest and the least material
damage was expected to be found.
Approach. Ring Sections 1 and 2 were visually
inspected and photographed and then were po-
lished and photographs were taken to document
the observed creep damage. Two portions from
Ring Section 1 and two portions from Ring Sec-
tion 2 were removed, mounted, polished, and
etched with glyceregia. Photomicrographs were
taken to document the creep voids prior to etching
and to document the microstructure after etching.
Results and Discussion. Figure 18 is a photo-
graph of Ring Section 1 following polishing. The
ring section was approximately 220 mm in diame-
ter by14.5-mm thick by 64 mm wide. Ring Sec-
tion 1 contained an inlet port for a pigtail, as indi-
cated in the figure. The port was approximately
31.8 mm in diameter and the wall was thicker at
this location, approximately 15.9-mm thick.
Figure 18. Photograph of Ring Section 1 after po-
lishing.
Indicated and visible on the polished surface is an
OD surface breaking crack that is within the limits
of the previously mentioned port. Although not
visible in the figure, a crack of similar size is lo-
cated on the ID surface of the port, below the visi-
ble crack.
Ring Section 2 was approximately 217 mm in di-
ameter by 14.3 mm thick. The location of interest
on the ring had a red marking.
Metallographic Examination. Figures 19 and 20
are photomicrographs of Met Section 1a (removed
from Ring Section 1) taken adjacent to the OD
and ID surface, respectively. Met Section 1a was
removed near the port of the pigtail inlet. The re-
sidual life of the tube section that Ring Section 1
was removed from is approximately 20% based on
the metallurgical analysis and guidelines of past
studies [33]. Based on the smaller size of the indi-
vidual voids in the figures compared to the thicker
width of the cracks indicated in Figure 19, it ap-
Port for
pigtail inlet
Cracks
Polished
surface
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2011 [96] AMMONIA TECHNICAL MANUAL
pears that the material tore after the creep voids
aligned. This tearing of the material suggests, but
does not indicate, that high-temperature creep was
involved. A high density of voids, but no cracking
was observed near the ID surface. The aligning
and linking up of the voids appears to follow the
dendritic columnar structure adjacent to the OD
surface. The location of the voids adjacent to the
ID surface suggests an equiaxed fine-grained
structure. The density of the voids is higher adja-
cent to the ID surface.
Figure 19. Light photomicrograph of Met Section
1a, taken adjacent to the OD surface and near the
pigtail inlet (as polished).
Figure 20. Light photomicrograph of Met Section
1a, taken adjacent to the ID surface and near the pig-
tail inlet (as polished).
Met Section 1b was removed from Ring Section
1opposite the port of the pigtail inlet. The density
of the voids throughout Met Section 1b was low.
Minor aligning of the creep voids was visible ad-
jacent to the OD surface.
Figure 21 is photomicrograph of Met Section 2a
(removed from Ring Section 2) taken adjacent to
the OD. The residual life of the tube section that
Ring Section 2 was removed from is approximate-
ly 60% based on the metallurgical analysis and
guidelines of past studies [33]. Met Section 2b
was removed opposite from the location marked
in red on the OD surface. The density of voids in
on both met sections, throughout the wall thick-
ness, is uniform and high. The voids adjacent to
the OD surface are aligned (as shown in Fig-
ure 21) and appear to follow a dendritic columnar
structure. The location of the voids adjacent to the
ID surface suggests an equiaxed fine-grained
structure. Thus, even though Ring Section 2 had
less creep damage than Ring Section 1, it still had
a significant amount of creep damage.
Figure 21. Light photomicrograph of Met Section
2a, taken adjacent to the OD surface and near the
red marking (as polished).
Figures 22 and 23 are photomicrographs of the
microstructure of Met Section 1a (removed from
Ring Section 1) near the pigtail inlet. The figures
show primary carbides located at the grain boun-
daries and secondary carbides dispersed in the
grains. The network of primary carbides is more
Cracks/aligned
creep voids
Aligned
creep voids
96 2011AMMONIA TECHNICAL MANUAL
2011 [97] AMMONIA TECHNICAL MANUAL
intact near the ID surface than near the OD surface
which indicates that the OD surface was subjected
to higher temperatures than the ID surface.
Figure 22. Light photomicrograph of Met Section
1a, taken adjacent to the OD surface and near the
pigtail inlet (glyceregia etchant).
Figure 23. Light photomicrograph of Met Section
1a, taken adjacent to the ID surface and near the pig-
tail inlet (glyceregia etchant).
Figure 24 is a photomicrograph of the microstruc-
ture of Met Section 1a (removed from Ring Sec-
tion 1) adjacent to the OD surface. A demarcation
is visible in the figure. There is no evidence of a
network of carbides above the demarcation and
there is a network below the demarcation. The re-
gion where the demarcation is present was likely
overheated since there is a severe reduction in the
amount of primary and secondary carbides there.
Figure 24. Light photomicrograph of Met Section
1a, taken adjacent to the OD surface and near the
pigtail inlet showing the decarburized region (glyce-
regia).
For the microstructure of Met Section 1a (re-
moved from Ring Section 1), opposite the port of
the pigtail inlet, the network of primary carbides is
less intact and more secondary carbides are dis-
persed in the matrix adjacent to the OD surface as
compared to the ID surface.
For the microstructure of Met Section 2a (re-
moved from Ring Section 2), at locations 180°
from each other, the network of primary carbides
adjacent to the OD surface at both locations is less
intact than at the ID surface. Again, the structure
of the primary carbides indicates that the OD sur-
face was subjected to higher temperatures than the
ID surface. The quantity of secondary carbides
appears to be similar when comparing the loca-
tions adjacent to the OD and ID surface.
Summary. A summary of the results follows:
• For Ring Section 1, the width of the creep
cracks near the port for the pigtail inlet suggests
a high temperature creep mechanism. A decar-
burized region near the port indicates that the
material was locally overheated. The creep
Demarcation
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2011 [98] AMMONIA TECHNICAL MANUAL
cracks and decarburized region were both lo-
cated adjacent to the OD surface. Aligned creep
voids, but no cracking, were identified in Ring
Section 2.
• The remaining lives of Ring Section 1 and 2
were estimated to be 20% and 60%, respective-
ly.
Creep-Rupture Strength
Jaske [34] has previously reviewed the creep-
rupture strength of weldments for hot outlet head-
ers and compared their stress-rupture behavior
with that of the cast 20Cr-32Ni-Nb alloy. Howev-
er, whereas as test data were available for the
weldments, only manufacturers’ curves (no data
points) were available for the cast base metal. The
lack of such published data makes it difficult to
assess the cause of the recent base metal failures
of hot outlet headers. Some operators and refor-
mer design engineers are of the opinion that the
creep-strength of the 20Cr-32Ni-Nb alloy is ac-
tually well below the published curves. For this
reason, they apply large design margins in estab-
lishing stress and temperature limits. Develop-
ment of reliable creep and stress-rupture data for
the 20Cr-32Ni-Nb alloy would be highly desira-
ble.
Conclusions
Hot outlet headers for steam-methane headers
have historically been subject to failure issues.
Early failures were related to the use of low-
strength alloys until the use of Alloy 800H or
800HT or 20Cr-32Ni-Nb became the construction
materials of choice. Then, more failure problems
became associated with welded joints and weld
metals with low creep strength. The strength and
quality of weld joints has increased, but they are
still prone to problems, especially when un-
planned for bending loads are induced in the
header. More recently, several failures have been
associated with the cast 20Cr-32Ni-Nb base metal
bringing into question its creep-rupture strength.
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