Volume 6 Preprint 69
In advanced Creep Failure of H.P. Modified Reformer Tubes in an Ammonia Plant
H. Shariat, A.H. Faraji, A. AshrafRiahy and M. M.Alipour
Keywords: reformer furnace, HP-modified Alloys, catalyst, overheating, creep
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Volume 6 Paper H012
In advanced Creep Failure of H.P. Modified
Reformer Tubes in an Ammonia Plant
M.H. Shariat1, A.H. Faraji2, A. AshrafRiahy3, M. M.Alipour4
of Materials Engineering, Shiraz University, P.O. BOX
7134851154, Zand Blvd., Shiraz, Iran, email@example.com
2General Engineering Department, Razi Petrochemical Co., P.O. BOX
161, Bandar Imam, Iran, firstname.lastname@example.org
3Department of Materials Science and Engineering, Tarbaiat Modarres
University, P.O. BOX 14115, Tehran, Iran, email@example.com
4Department of Materials Engineering, Shiraz University, P.O. BOX
7134851154, Zand Blvd., Shiraz, Iran, firstname.lastname@example.org
Ammonia plant in Razi petrochemical complex has been recently
revamped by M.W. Kellogg. This work involved the replacement of HK40 tubes with HP Modified (Manaurite XM) supplied by Manoir
industries, France. After 18 months of operation, one of the tubes
failed. Because of in advanced failure, an investigation to find the
causes of this failure started in joint effort with Tarbiat Modarres
University and Shiraz University, Metallurgical studies showed that
tube failure and creep occurred as a result of excessive temperature
due to chocking in this particular tube. In this paper the results of this
investigation is presented.
Keywords: reformer furnace, HP-modified Alloys, catalyst, overheating,
Steam Methane Reformer Furnaces
The steam-gas reformer is a common but critical piece of equipment
in ammonia and methanol plants. In most plants, methane is used as
feedstock. It reacts with steam in catalyst-packed tubes at high
temperature. The process is highly endothermic. The tubes have inside
diameters of 60-120 mm (2.5 in.) and are 10-14 m (33-46 ft.) long.
The pressure is 15-30 bar (218-435 psi) and the temperature between
900 to 1000°C (1652-1832°F). The tube wall thickness ranges from 8
to 20 mm (0.31-0.79 in.) depending on the tube diameter,
temperature and pressure. Excess steam is used to reduce the
formation of carbon. The reforming reactions are favored by high
temperature but retarded by high pressure. A typical steam methane
reformer for hydrogen and methanol production is shown in Figure 1.
When synthesis gas is used in the production of ammonia the
conditions are considerably different. Nitrogen is provided by the
injection of air into the primary reformer effluent and a secondary
reforming step is carried out in a separate catalyst- filled vessel.
Because of the balanced load between the two reformers, the
temperature in the primary reformer tends to be lower, 815°C (1500°F)
and the pressure between 15 and 40 bar (218-580 psi) .
HP Modified Alloys
Since the catalyst tubes assembly can amount to 25% of the total cost
of the furnace, there is a great incentive to optimize the design
chemically. There are many approaches in satisfying the process and
operational requirements. One of the most dramatic improvements in
reformer cost has been the increased application of stronger,
proprietary tube materials such as HP-Mod alloys. In new reformer
design, the HP-Mod alloys have allowed the size of the reformer tube
to increase, particularly those designed for the high-pressure
ammonia process. With the HK-40 alloy the reformer tube diameter
was limited by the tube wall thickness. Excessively thick tubes are
candidates for accelerated creep damage resulting from large
temperature differences across the thick tube wall. The use of HP-40
Mod Nb has allowed the internal volume of tubes to grow dramatically,
without increase in the tube wall thickness, producing capacity
increase of 30% to 40% for the same tube count, with only marginally
Basically, the alloys used for high temperature purposes rely upon
creep strengthening by the formation of carbides in the microstructure
(Figure 2). There are two main forms of carbide:
Primary Carbides, which form as a grain boundary network
during solidification (Figure 2.a).
Secondary Carbides, which form as fine particles within grains
during service (Figure 2.b).
The secondary carbides are the main contributors to creep strength.
Within just the first few hours of service, the microstructure of new
tubes will change as secondary carbides precipitate inside the
austenitic grains (a process known as “ageing”). Each small carbide
particle acts as a barrier to prevent the movement of creep damage
through the microstructure. Therefore, the finer this dispersion of
secondary carbides, the greater the creep strength. As service time
and/or service temperatures increases, the secondary carbides diffuse
and become coarser. With the loss of secondary carbides from the
austenite matrix, the creep strength is reduced.
Whilst this loss of secondary carbides occurs slowly at normal
operating temperatures (below 950°C (1742°F)), it occurs more quickly
at higher temperatures, thus overheating for just a short time can be
so damaging to tube life. This sensitivity to overheating is most
common with the older type alloys such as “HK 40”. Alloy development
has therefore concentrated upon producing alloy grades that have
greater resistance to loss of secondary carbides. Addition of alloying
elements like nickel and niobium in new grades ensure that the
secondary carbides were finer and more stable for longer times at
elevated temperatures and thus have greater creep strength than the
more simple, Nb-free “HK 40” which was the first to be developed 40
years ago (Table 1).
As mentioned earlier, the improved metallurgy of the new alloy grades
allow steam reforming furnaces to operate at increasingly higher levels
of temperature, pressure and heat flux. In general, the advantages of
using a micro alloyed grade such as HP-40 Mod Nb are:
Lower tube skin temperatures
Thinner tube walls
Lower thermal stress
More efficient fuel utilization
This is because the higher strength material requires a smaller
minimum wall thickness for the same operating conditions. Tube wall
thickness is normally calculated using the design standard API RP530.
This recommends that the allowable stress should be the statistical
minimum stress (the lower line of normal scatter band). Once the
design temperature is known, it is possible to define the mean
100,000 hour stress at that temperature. Some engineers will use a
percentage of the mean stress (say 80% of the mean value) as the
allowable stress. This three dimensional study of
temperature/stress/life is, however, a complex matter and usually
computer aided statistical study based on the Larson-Miller
extrapolation is used to give the true minimum value of sound wall.
The definition of temperature is important. The gas temperature is the
normal basis for controlling the process (catalyst and feedstock). It is,
however, the tube skin temperature that is the basis for the design
performance of the tubes themselves. The hotter the tube, the lower
the creep strength and vice versa. These can be significantly different
depending on the tube thickness and furnace operation. In practice,
the life of the tube not only depends on the skin temperature but also
the stresses imposed on the tube.
The most important property of a reformer tube alloy is the 100,000
hour Creep Rupture Strength which is defined as “The stress which
alloy can withstand for a lifetime of 100,000 hours at design
temperature before creep failure”. As stronger alloys have been
developed, this resulted in a large increase in the 100,000 hour Creep
Rupture values .
Common Damage Mechanisms in Reformer Tubes
It is well known that the main damage mechanism, which reduces tube
life, is creep. This occurs in metals under stress at high temperature
and is designed into the operation. If the recommended conditions are
followed, creep occurs slowly over a period of about 11 years
(100,000) or more. If, however, the tubes are overheated the tube life
will be affected. This does not generally result in immediate failure but
will certainly result in a reduction in life. Depending on the degree of
overheating the reduction can be dramatic. This has been displayed
extremely well by the data in Table 2, which shows the dramatic effect
of the logarithmic relationship between temperature and life according
to API R530.
It can be seen that quite small increase in tube temperature over a
long period of time will still reduce the life significantly. This is often
not appreciated when temperatures are increased to maintain yields
when catalyst becomes less active. The much greater effect, however,
is the rapid increases which occur during feedstock or steam supply
failures. The absence of the cooling effect of the endothermic reaction
can result in very high temperatures which will give a very short
theoretical life. At temperatures around 1235°C, incipient melting of
the carbide eutectic occurs and the material becomes plastic. It is also
important to realize that the effects of overheating are cumulative.
Once creep life is lost it can not be regenerated. If, however, the tubes
are operated below design conditions, the remaining life will be
extended through slower creep rate.
This is a further contributing factor to accelerated creep. When tubes
are heated and cooled, through wall stresses are temporarily increased
and consequently a proportion of life is lost. Tube wall temperature
gradients can be significant. During normal operation the stress
produced by different expansion relaxes through creep. Temperature
changes up or down will reintroduce some stress.
This is another relatively common problem. Heating to operating
temperature results in linear expansion, which must be accommodated
in the furnace design. Some older designs of furnace are known to be
susceptible as a result of inadequate or non-existing top tensioning
which results in an effective restraint at the furnace roof. Modern
furnaces are generally much better designed. With tensioning based on
about 80% of tube weight, vertical expansion is encouraged without
unduly increasing the stress. In some designs, a spring system will
help prevent stress to be created as a result of bending.
Thermal shock creates extremely high stresses as the tube attempts to
contract under restraint. This will result in circumferential tearing and
in extreme cases, shattering of the tube. This has been shown to occur
through boiler water carry over on the inside of hot tubes or rainwater
on the outside above the furnace roof.
Stress Corrosion Cracking
Some designs of furnace having a cooler dead space at the tube top
and/or bottom can be susceptible to stress corrosion cracking. If
steam is allowed to condense and re-evaporated in these areas there
can be significant amount of impurities which are concentrated by the
evaporation. This can be particularly severe in designs having
refractory in contact with the steam .
Razi Reformer Tube Failure
The Ammonia plant at Razi Petrochemical Company has been recently
revamped by M. W. Kellogg Ltd. This work involved replacement of the
existing centrifugally cast HK-40 reformer tubes with new HP-
Modified ones (Manaurite XM), supplied by Manoir Industries, France.
There was a failure of one tube after 16 months operation.
Subsequently samples of this tube were collected from various
locations including the failure position and metallurgical investigation
was done to make an assessment of the failure mechanism .
Reformer and Catalyst Tube Details
The reformer is a top fired design with 360 vertical, top hung tubes.
Design calculations advise that the hottest position on the tubes is
near the base. A schematic of the heater showing the layout of the
radiant section of the tube is shown in Figure 3. In operation, the
tubes are filled with a nickel based ring catalyst that reforms the
natural gas/steam feed at a pressure about 31-33 bar(g).
Three Manaurite XM spool pieces are joined to make the bottom
section of each reformer tube, the top section being fabricated from
A335 – Gr. P11. A number of such tubes are welded directly to an
outlet header to form a harp. These harps, prefabricated at Manoir, are
shipped to site where they are assembled into tube rows .
Samples and Visual Examination
The split are of the tube showed a large local bulging. A rough
measure of tube circumference in the bulged area showed a 350 mm
value compared to 337 mm in adjacent visually non damaged area.
Nominal tube OD is 104.8 (+0.2) mm which gives a nominal
circumference of 329 to 335 mm. Visually the failure has a typical
creep aspect with numerous longitudinally oriented cracks. Obviously
it is not a brittle failure (Figures 4,5).
Five samples labeled A, B, C, D1 and D2 (Figure 6) were collected for
evaluation. The locations from which the samples were taken are
shown in Figure 7. Visual examination of samples A and B did not
reveal anything untoward. Examination of samples C and D2 revealed a
smooth inner surface to the tube typical of a machined finish. Sample
D1 showed undulations oriented in the axial direction typical of
macrostrain accumulation (Figure 8). The outer surfaces of samples D1
and D2 appeared unusually smooth for centicast tubes; however,
discussions with manufacturer revealed that the failure occurred within
the area machined as part of the manufacturing process for this
section of tube.
Analysis and Results
The analysis and results of assessment are given below:
Chemical Analysis of Sample D2. The tubes have been fabricated
according to Manoir Industries Manaurite XM specification. Details of
the specification together with the results of the analysis on a section
of the tube from D2 are given in Table 3. From the table it can be seen
that the analysis of the tube sample lies within the manufacturer’s
specification. The elements Nb, Ti, Zr and others will have a significant
effect on creep strength of the material, and this analysis indicates
that these elements are present. However it is not possible to compare
the sample D2 microalloy content against any detailed specification,
since this is Manoir’s proprietary information.
Minimum Sound Wall for each Sample. Through wall thickness
measurements were taken on all samples using a traveling
microscope, one measurements were taken across each sample A and
B, measurements were taken every centimeter around samples C, D1
and D2. The results are given below.
Examination of all samples showed an absence of any casting porosity
or other bore related manufacturing defect. Thus measured wall
thickness can be taken as equivalent to actual sound wall.
Samples A, B, C, and D2 are slightly above specification with regard to
minimum sound wall thickness. Sample D1, which is adjacent to the
failure, shows reading lower than the other tubes and within
specification. This is almost certainly a result of the strain
accumulation associated with the failure, rather than a difference in
original wall thickness.
Metallographic Examination of all samples and the failure mode of
sample D1. All etchings were carried out electrolytically with 5% oxalic
acid. The samples were examined in the radial plane, at right angles to
the orientation of cracking. Comments on the findings of the
microstructural examination are given below.
Examination of the through wall samples revealed typical centricast
structures with columnar grains adjacent to the outer surface and
equiaxed grains near to the inner surface.
Sample A – A network of primary carbides was apparent on the grain
boundaries, consistent in appearance with an HP Modified furnace tube
on entering service (Figure 9).
Sample B – The network of primary carbides was still clearly apparent.
A fine dispersion of secondary carbides has precipitated within the
grains (Figure 10).
Sample C – Examination of this sample revealed large primary carbides
on the grain boundaries and large secondary carbides within the
grains. Creep cavities were also apparent on the grain boundaries on
this sample (Figure 11).
Sample D2 - Examination of this sample revealed large primary
carbides on the grain boundaries and large secondary carbides within
the grains. Microcracking was apparent in the outer third of the tube
wall oriented in the through wall direction; oriented cavitation was
apparent through the sample (Figure 12).
Sample D1 - Examination of this sample revealed large primary
carbides on the grain boundaries and large secondary carbides within
the grains. Macrocracking, microcracking and cavitation were apparent
through the tube wall oriented in the through wall direction. There was
heavy precipitation of secondary carbides associated with some of the
larger cracks. Again the highest concentration of cracking was seen in
the outer third of the tube wall (Figure 13).
Both the location of the damage in the tube wall and its appearance
are typical of creep damage resulting from high temperature operation
under an internal pressure stress. The lower levels of microstructural
degradation in samples A and B, when compared with samples C, D1
and D2, indicate that damage is unlikely to occur at these locations in
the short term.
Estimation of Temperature Inducing Creep Rupture after 13,000 hours
in the Tube at Design Pressure
Design Pressure: P = 32 bar(g) = 3.2 MPa
OD 105 mm
Thickness t = 12.5 mm
Stress in tube section = σ :
P × OD P 3.2 × 105 3.2
2 2 × 12.5
σ = 13.44 - 1.6 = 11.84 MPa
According to XM Larsen Miller curve in Figure 14:
σ = 11.84 MPa
→ LMP = 35.80
35.8 = T[22.96 + log(13,000)] × 10-3
35800 = T(22.96 + 4.114)
T = 1322 K = 1049°C
Remarks on the Observed Failure
The examination carried out here indicates that the tube failed as a
result of the application of the operating pressure stress under
excessive temperatures. The tube composition is consistent with the
manufacturer’s specification and the tube microstructure in sample A
is consistent in appearance with an HP Modified alloy in its virgin state.
Possible causes of overheating in the reformer furnace tube are listed
1. Burner misalignment resulting in over temperature tubes.
2. Restricted flow of process through tube due to catalyst choking.
1. The tube failure occurred as a result of excessive operating
temperatures (1049°C). Samples C, D1 and D2 all contain
significant creep damage, distributed in a manner indicative of
pressure stress loading under overheat conditions rather than
through wall thermal stress loading under normal conditions.
2. Chemical analysis revealed sample D2 to be consistent in
composition with Manaurite XM.
3. Thickness Measurements indicate that all the tube walls
measured, apart from that immediately adjacent to the failure,
are above maximum specified in the design.
4. Samples A and B, taken from higher positions in the tube,
represent material appropriate for further extensive service.
They show no significant creep damage.
1. C. M. Schillmoller, HP-Modified Furnace Tubes for Steam
Reformers and Steam Crackers , NiDi Technical Series No 10 058,
2. Manaurite 40 X Product info, Manoir Industries, France.
3. David Pool and John J. Jones, Metallurgy and Material Selection
for Steam Reforming Furnace in Ammonia & Urea Plants,
DONCASTERS Industries, United Kingdom, 1999
4. G. Verdier, Catalyst Tube Failure, Report No. 1376, Manoir
Industries, France, 2000.
5. T. Lant, Metallurgical Investigation of Reformer Tube Failure,
ERA Technology, United Kingdom, 2001.
Table 1. Nominal composition of Principal Reformer Tube
“HP Nb Mod.”
“IN 519 TZ”
Table 2. Decrease in tube life with increase in temperature for HP
Nb Mod. Alloy
2 ½ years
4 ½ months
5 ½ days
Table 3. chemical Analysis of Sample D2, comparison with Manaurite XM Spec.
Nb, Ti, Zr and others
Figure 1. Steam methane reformer, typical for hydrogen production
Figure 2. (a) Primary Carbides, (b) Primary and Secondary Carbides
Figure 3. The Radiant Section of Reformer Furnace
Figure 4. Local Bulging at the Split Area of the Tube
Figure 5. Longitudinal Cracks at the Bulged Area of the Tube
Figure 6. Samples Collected From the Failed Tube
Figure 7. Sketch Showing sample locations from catalyst tube
Figure 8. Inner surface of sample D1 with strain lines apparent
Figure 9. Sample A with primary carbides clearly visible
Figure 10. Sample B with Primary and Secondary Carbides
Figure 11. Primary and Secondary Carbides and oriented cavities in
Figure 12. Microcracking and oriented cavitation in Sample D2
Figure 13. Sample D1 showing Macrocracks
Figure 14.Lason-Miller Curve for Manaurite XM