Cracking of Alloy 800H Reformer Riser - Quest … 11 - Cracking of Alloy 800H Reformer Riser CHARLES THOMAS, QUINTON ROWSON Quest Integrity Group NZL Limited Lower Hutt, New Zealand

- 11 -

Cracking of Alloy 800H Reformer

Riser

CHARLES THOMAS, QUINTON ROWSON

Quest Integrity Group NZL Limited

Lower Hutt, New Zealand

A riser in a Kellogg design reformer failed. This failure was attributed to water quenching of the riser tube just below the weld between the cast HP alloy riser tube and an alloy 800H transition piece. The water originated from a leak in the water jacket. During the investigation however, additional cracks were found in the transition pieces. This cracking could not be attributed to the quenching incident. This additional cracking was associated with voiding and oxidation along the fracture and was diagnosed as a creep failure and repaired. Finite element analysis of the riser design ruled out system loading as the source of the stress that caused the cracking. The repaired riser was subsequently found to have re-cracked along with other risers in the reformer. A root cause analysis exercise was undertaken to review all of the contributory factors that led to the failure. The cracking was ultimately attributed to stress relaxation cracking and heat treatment was undertaken to mitigate this damage mechanism. Stress relaxation cracking of austenitic alloys has occurred in a number of plants in recent years and its occurrence has resulted in questions being raised about the generally held belief that austenitic alloys do not require post weld heat treatment. This paper discusses the implications of this on steam reformers.

1. INTRODUCTION The Kellogg reformer design is unique. The reformed gas from the collection manifold at the bottom of each row of reformer tubes is transferred through a riser, returning back up the centre of each row to the effluent chamber at the top of the furnace. The arrangement is shown schematically in Figure 1. The process gas effluent passes up the riser, through the furnace roof and into a refractory lined transition piece that connects to the effluent chamber. The transition piece and effluent chamber are cooled by a water jacket. Within the furnace, the riser tubes operate at a temperature controlled by the combustion environment within the furnace and that of the process stream within the tube. The metal temperature is of the order of 780°C. On leaving the furnace through the roof, the external metal temperature reduces rapidly, over a short length, to that of the water jacket.

C. Thomas, Q. Rowson

12 Nitrogen + Syngas 2014 International Conference & Exhibition (Paris 24-27 February 2014)

In this plant, the risers had been replaced a number of times over the full history of the plant. The most recent replacement exercise took place in 2006 and was accompanied by a design change. In 2006, the riser diameter was increased to accommodate an increase in production rates. At this time, the material of construction for all the risers was changed to the micro alloyed version of 25Cr35Ni HP alloy (HPMA) which is commonly used in catalyst tube applications. The original design had a material change from the cast alloy within the furnace to a wrought alloy 800H forging just below the water jacket. This material change in the original design was located within the refractory close to the furnace roof. The opportunity was taken to raise this transition weld above the furnace roof by increasing the length of the cast riser from 11.1 metres to 11.4 metres, lifting the weld clear of the furnace refractory, to facilitate the making of the site connection weld and subsequent inspection. The arrangement is shown in Figure 2. Above the furnace roof, the short length of cast HPMA material was contained within an insulation can that ended very close to the forged alloy 800H fitting. It was this connection between the cast HPMA riser and the alloy 800H fitting that had failed and cracked on a number of occasions.

2. THE FAILURE In early 2010, the plant was taken off-line to investigate large steam clouds that were forming around the riser on Row #4 of the reformer above the furnace roof. The steam had been caused by water leaking from the water jacket around the effluent chamber above riser #4 and running down the riser. On contact with the main riser tube below the water jacket, the water flashed to steam. Inspection of the riser just below the weld to the alloy 800H forging found a large circumferential crack. The area of cracking in the riser coincided with staining on the water jacket indicating the location down which water had flowed.

A ring was removed from the top of the riser to confirm the failure mechanism. The crack that was found is shown in Figure 3. The HP alloy section of riser had very low toughness; Charpy impact testing at room temperature provided Charpy fracture energy of only 4 Joules. It was concluded the failure of this component, which led to a through wall crack, was indeed due to the thermal stress caused by the quenching effect of the water leak from the water jacket. The location of this crack and the staining from the water leak were coincidental. The features of the cracks formed in service were similar to brittle cracks generated in the laboratory and are consistent with brittle failure. Specifically, the cracks were constrained to the interdendritic primary carbides and it could be seen that these carbides had appeared to “shatter” consistent with a single incident brittle failure. The low toughness of this material as revealed by the Charpy impact tests suggests that this material would be susceptible to brittle failure. This low toughness was further demonstrated by problems encountered during welding and the need for solution anneal heat treatments.

The background microstructure of the HPMA cast riser tube was as expected for an ex-service HP material. The absence of creep damage and the modest levels of microstructural change caused by exposure to service temperature suggested that prior to the brittle failure, this component was in good condition.

It was concluded therefore that the failure of the HPMA riser transition was due to the water quench and that in the absence of similar thermal shocks, other HPMA riser components were likely to be in good condition. It should be noted however, that the low toughness of the failed HPMA riser at low temperatures is “normal”. Other HPNb and HPMA components are likely to have similarly low levels of toughness and are similarly prone to thermal shock, either by water quench as in this case, or mechanical impact. Given the poor toughness, similar components should be guarded from thermal shock or mechanical impact.

The material of the riser tube was analysed. It was found to be in accordance with the specification for the HPMA alloy. This cracking was considered to be consistent with the quenching of the cast riser tube by water leaking from the water jacket as expected from the circumstances of the failure.

At the same time however, while examining the ring sample taken from the top of the riser, a second cracking mechanism was observed in the heat affected zone of the weld, between the top of the spun cast riser tube and the alloy 800H forged transition piece. This crack is illustrated in cross-section in Figure 4. The crack was located in the heat affected zone within the alloy 800H transition cone forging at the top of the cast section of the riser. It is of very different nature to the brittle crack in the HPMA cast alloy. The

Cracking of Alloy 800 H Reformer Riser

Nitrogen + Syngas 2014 International Conference & Exhibition (Paris 24-27 February 2014) 13

crack was parallel to the weld fusion line, contained significant amounts of corrosion product and was associated with numerous voids. The conclusion reached was that this was a creep or creep fatigue crack which implied significant axial loading on the riser.

The riser was repaired and returned to service. The following year, inspection found all 4 risers to be cracked in the heat affected zone of the alloy 800H transition piece. These were repaired. Following a further failure in 2012, the decision to undertake a formal root cause analysis was made.

3. ROOT CAUSE ANALYSIS The plant owners assembled a team of stakeholders and experts including:

Plant Manager

The lead Asset Manager

The organisation’s global integrity pressure systems leader

Senior process coordinator

Welding Supervisor

Mechanical Engineer

Metallurgical Engineer.

This team was led by a root cause analysis expert. Following extensive discussion to identify all possible contributory factors no matter how likely or unlikely, a formal Causal Tree was produced. This identified in excess of 50 questions to be answered. These were classified by cost and importance. The tasks allocated to metallurgical and fitness for service assessments are listed in Table 1.

Table 1 - Potential causes of cracking and solutions identified in the RCA and

allocated to Quest Integrity

Actions Cause Solution

1 Relative Location of Weld, Water jacket and Insulation

Review whether or not relaxation cracking is a credible riser failure mechanism

2 Review whether changes to welding procedure to reduce likelihood of relaxation cracking at the weld location are required. - Check if Post Weld Heat Treatment is an option

3 Bottom Header restricted Movement

Review the impact on the bottom header movement due to the 2006 Modification of the bottom header

4 Unexpected Movement of effluent chamber

Conduct a stress analysis on movement of the effluent chamber and determine what effect it can have on the risers

5 Poor Design of 800H components Develop a model of the Riser to effluent chamber tubes to determine thermal and pressure stresses

6 Variation of Feed Rate Model the impact of Plant Upsets, Plant Trips and Plant Cycling on the riser tubes

4. STRESS RELAXATION CRACKING A review of the phenomenon of stress relaxation cracking was undertaken to address action items 1 and 2. The phenomenon of stress relaxation cracking in austenitic materials has been researched by the TNO organisation in the Netherlands led by Hans Van Wortel [1]. It would appear that cracking of welded austenitic components when operated (or post weld heat treated) at elevated temperatures has occurred in a range of industries and is known by a wide range of terms including:

Relaxation Cracking

Stress Induced Cracking

Reheat Cracking

Stress Relief Cracking

Post Weld Heat Treatment Cracking

Strain Age Cracking

White Phase Fractures.



These terms all refer to a phenomenon in which the response to stress occurs by grain boundary cracking as opposed to plastic deformation (i.e. creep) due to the local metallurgical condition. This typically occurs when the individual grains are strong and resist creep deformation while the grain boundaries are relatively weak. In the case of creep-resistant ferritic steels such as P11 or P22, this may occur in the coarse grained heat affected zones of welds. These coarse grained areas are relatively creep-strong and the response to stress is cracking of grain boundaries as opposed to plastic deformation (creep) of the grains.

Van Wortel’s paper [1] refers specifically to the cracking of austenitic alloys such as 321 and 347 stainless steels, nickel based alloys such as 617 and of most relevance in the present case, alloy 800H. He categorizes typical features of cracking in austenitic materials including alloy 800H as:

The cracks are always located on the grain boundaries and in front of the cracks, small isolated cavities are present.

The cracks are present in the Heat Affected Zone, weld metal or cold deformed areas.

Mostly, a metallic filament is present on the cracked grain boundaries. This filament is enclosed by a chromium-rich oxide layer. In this oxide layer, the Ni and Fe contents are low. The chemical composition of the metallic filament is material dependent but always low in chromium and high in nickel and iron.

The cracks are present in areas where the Vickers hardness is higher than 200HV.

The operating metal temperature is always between 550°C and 750°C and varies within this range for different materials. On average, the most critical area is around 600°C.

Furthermore, he states that most stress relaxation cracking failures occur within a year or so of service, which is consistent with the cracking and failures between 2010 and 2012. Alloy 800H contains small amounts of deliberately alloyed titanium and aluminium. These are common alloying elements in superalloys that obtain their strength from the formation of γ’ (gamma prime - Ni3(Al,Ti)) precipitates. These have been observed to form in alloy 800H and influence the mechanical properties, increasing strength and significantly reducing creep ductility [2]. The temperature at which the greatest effect was observed was 650°C. Above approximately 850°C, the precipitates coarsened and re-dissolved and their influence was significantly diminished. This effect provides a mechanism by which the alloy 800H may strengthen and become resistant to relaxation by creep. This precipitation would also be expected to cause an increase in hardness and is consistent, for alloy 800H at least, with the observation by van Wortel [1] that relaxation cracking occurs in material with hardness above 200HV. Van Wortel, however, indicated that the hardening in alloy 800H was due to fine precipitation of chromium carbides (Cr23C6).

Notwithstanding the hardening by γ’ or carbide precipitation, grain growth in the heat affected zone will also increase creep strength and reduce ductility making the material more susceptible to relaxation cracking [3]. Alloy 800H is deliberately heat treated to increase grain size above ASTM grain size #5. Coarse grain size materials are more prone to stress relaxation cracking than fine gain size materials.

A pre-requisite for oxidation resistance in the majority of alloys used at elevated temperature is chromium; alloy 800H contains approximately 20%. This is present to provide a thin but continuous surface layer of oxide that protects the material from further attack. A side effect of the presence of chromium is the formation of internal chromium carbides most commonly Cr23C6. These carbides form predominantly on grain boundaries and the localised removal of chromium from solution in the grain has the effect of weakening the area adjacent to grain boundaries. This is the reason why cracking frequently is associated with the formation of a metallic filament. The weakened material either side of the grain boundary cracks, leaving the original material along the grain boundary [3]. When exposed to an external environment, the chromium-depleted area is also reduced in resistance to oxidation and an oxide layer is formed.

The riser design was modified in 2006. Prior to this design change, no cracking at the top of the risers had occurred. The design change involved moving the welded joint between the top of the riser tube and the alloy 800H transition piece higher above the furnace roof (Figure 2). Prior to 2006, this weld had been inaccessible and lay within the insulation at the furnace roof. This change would reduce the temperature at which the weld was operating. Prior to 2006, there would have been little or no cooling of the reformed gas



exiting the furnace by the time it reached the alloy 800H transition piece, meaning it would operate at approximately 780°C. After 2006, the weld was relocated approximately 500 mm above the furnace roof, close to the top of the insulation can and close to the bottom of the water jacket. This would result in a significant reduction in the operating temperature of the cracked alloy transition piece (Figure 2).

The circumstances for stress relaxation cracking therefore exist within the redesigned riser. A review of the original metallurgical analysis showed that many of the features of stress relaxation cracking were present (Figure 4), including:

The crack was intergranular and within the heat affected zone.

The crack was associated with voids ahead of and adjacent to the crack.

The crack had a white filament within an oxide. The wider open areas of the crack had an oxide layer with the remains of what could have been a metallic filament on the surface of the oxide.

The filament was nickel and iron rich and chromium depleted while the oxide was chromium rich; the results of analyses of these zones using an Energy Dispersive Analysis technique are shown in Table 2.

Table 2 - Analysis using Energy Dispersive Analysis (EDS) system and a scanning

electron microscope

%O %Al %Si %Ti %Cr %Mn %Fe %Ni

Filament 0.5 0.3 0.47 13.7 1.16 50.1 33.77

Oxide 18.05 0.68 0.59 0.58 37.64 2.04 24.01 16.41

Bulk

Alloy 0.47 0.53 0.38 19.44 1.04 46.7 31.43

Alloy 800H 0.15-0.6 1.0 max 0.15-0.6 19-23 1.5 max Bal 30-35

In this case, the requirement for stress relaxation cracking, that the hardness be in excess of 200HV, was not met. The hardness was measured on the available sample to be 160-195HV in the vicinity of the crack. However, in the sample that was available for hardness testing, the primary failure mechanism had been quenching of the HPMA riser tube. The possible stress relaxation cracking had not proceeded to complete failure.

It was concluded following this review that, despite the low measured hardness, stress relaxation cracking was a credible damage mechanism leading to the cracking of the alloy 800H forging at the top of the riser tube.

5. GLOBAL STRESS ANALYSIS The solutions 3 to 6 as in Table 1 were all addressed by a single Finite Element Analysis model. The solutions listed in Table 1 make reference to specific locations within the system. However, in order to understand the global loading on the cracked weld at the top of the riser, the movement of the system as a whole was modelled from the mixed feed crossover line to the inlet of the secondary reformer. A global model was created that would determine loading at any chosen location but specifically for the purposes of the root cause analysis, the loading at the cracked weld.



The global finite element geometry for the model is shown in Figure 5. The boundary conditions selected as fixed points relative to which all displacements were assessed were:

Start of cross over line (inlet to primary reformer).

Connection between the effluent chamber and the secondary reformer (inlet to secondary reformer).

To establish relative movement and hence loading at all locations in the system, all factors were taken into account including:

Operating conditions, pressure and temperature

Component geometry and weight

Material properties

Thermal movement

All hanger settings and hanger loading

Examples of output from the analysis are illustrated in Figures 6 and 7. In these figures, the displacements have been exaggerated for illustrative purposes. Nevertheless, it can be seen that unexpected movements are predicted. The effluent chamber was predicted to tilt downwards somewhat above the reformer relative to the connection to the secondary reformer and this was accompanied by bending of the risers (Figure 6). The model predicted that the outlet headers at the bottom of each row of catalyst tubes would adopt a bow with the extreme ends moving upwards relative to the riser connection (bull-tee) in the centre of the manifolds.

With respect to the risers at the location of the cracking, the model predicted loading and un-relaxed elastic stresses as shown in Table 3. The predicted stresses are very low with the maximum occurring in Riser #1 at 6.8 MPa. It is noteworthy that the cracking and initial failure occurred in Riser #4 which was predicted to have a maximum stress of only 4.1 MPa at the failure location.

With respect to the action items identified in Table 1, the movement of the headers was determined (action item 3) and the model predicted these headers would adopt an upward bow. The implication is that this would place a compressive load on the risers, and indeed, this was the case. This compressive load was however, exceeded by the axial tensile load due to internal operating pressure. The global stress analysis suggested a greater concern should be applied to the design and loading of the bull tee to manifold connections than the risers.

Table 3: Summary of axial loading in risers from FEA model

Component

Axial load at joint* (KN)

Due to Gravity Due to Gravity +

Temperature Due to Gravity + Temperature

+Pressure (Operating conditions) Max Elastic stress MPa

Riser 1 2.1 (tensile) 14.87 (tensile) 68.45 (tensile) 6.8

Riser 2 -3.2 (compressive)

5.42 (tensile) 58.99 (tensile) 5.8

Riser 3 -8.1 (compression)

-3.69 (compression) 49.86 (tensile) 4.9

Riser 4 -12.7 (compression)

-12.58 (compression) 40.95 (tensile) 4.1

The analysis showed that under the current hanger configuration, it was unlikely that there would be any unexpected movement of the effluent chamber as it was fully supported by the effluent chamber hangers in the cold and operation mode. The finite element analysis showed that unexpected movement of the effluent chamber was not a factor leading to the damage observed in the risers (action item 4, Table 1).



The model also addressed action item 5, Table 1, and showed that thermal movement loads were relatively small compared to those caused by pressure. With regard to the impact of plant start-up and shut down cycles, these were not explicitly modelled. However, given the relatively low maximum load predicted in the riser it was concluded that load cycling was very unlikely to be a factor leading to the cracking observed.

In summary, the global analysis was successfully concluded and it demonstrated that loading on the risers, either from mechanical or system loads, was not the cause of the damage observed in the risers. In addition, the model has now become a valuable resource in understanding loading throughout the reformer and has identified potential issues concerning the bull-tee attachment welds to the manifolds. In this respect, the model has become a valuable contribution to the risk assessment of the plant.

6. CONCLUSION Based on the review of the failure, the phenomenon of stress relaxation cracking and the global finite element stress analysis, it was concluded that stress relaxation cracking was a significant contributor to the cracking and failures that had been observed. Mitigating damage due to relaxation cracking is described by van Wortel [1] and Shoemaker et al. [4]. The recommendation [4] is to apply a stress relief heat treatment at a minimum of 885°C for 1.5 hour plus 1 hour per inch of section thickness. Indeed, this heat treatment is a requirement of ASME Section VIII, Division 1 [5] specifically for alloy 800, 800H and 800HT.

It is of interest that it does not appear to be a requirement in the process piping code B31.3. Van Wortel [1] stated that in the programme of testing at TNO, as welded alloy 800H was extremely susceptible to stress relaxation cracking but post welded heat treated welds were not. The benefits of post weld heat treatment are potentially two-fold. Firstly, it will eliminate the weld residual stress. Secondly, it will reduce hardness and “over-age” the material reducing the potential for hardening due to γ’ precipitation. In the present case, post weld heat treatment is expected to achieve these benefits. A detailed description of the problem and the appropriate actions to mitigate stress relaxation cracking are to be incorporated into an API recommended practice currently in draft form [6].

This post weld heat treatment has been applied to the risers in question and no in service failures have occurred since. The results of the inspection to be undertaken at the next scheduled outage are eagerly awaited.

References 1. Hans van Wortel, “Control of Relaxation Cracking in Austenitic High Temperature Components”, Paper No. 07423,

Corrosion 2007, Nashville Tennessee, March 11-15 2007, NACE.

2. George Lai, “High Temperature Corrosion and Materials Applications”, Chapter 14 “Stress Assisted Corrosion and Cracking” page 398 published by ASM.

3. Christian E. van der Westhuizen, “Stress Relaxation Cracking of Welded Joints in Thick Sections of a TP347 Stabilized Grade of Stainless Steel. Paper No. 08454, Corrosion 2008, New Orleans, March 16-20 2008, NACE.

4. Lewis Shoemaker, Gaylord Smith, Brian Baker, Jon Poole, “Fabricating Nickel Alloys to Avoid Stress Relaxation Cracking”, Paper No. 07421, Corrosion 2007, Nashville Tennessee, March 11-15 2007, NACE.

5. ASME Boiler and Pressure Vessel Code, Section VIII, Division 1 rule UNF-56.



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Fig. 6: Vertical displacements looking in direction normal to the effluent chamber



Fig. 7: Vertical displacements looking in direction normal to the effluent chamber and normal to the outlet headers

Documents

Cracking of Alloy 800H Reformer Riser - Quest … 11 - Cracking of Alloy 800H Reformer Riser CHARLES THOMAS, QUINTON ROWSON Quest Integrity Group NZL Limited Lower Hutt, New Zealand