Superior Spun Cast Material for Steam Reformer Furnaces: Alloy

Centralloy®4852 Micro R

Dr. Dietlinde Jakobi Schmidt + Clemens GmbH + Co. KG, 51779 Lindlar, Germany

Dr. Peter Karduck and Prof. Dr. Alexander von Richthofen HDZ Herzogenrather Dienstleistungszentrum GbR, 52134 Herzogenrath, Germany

Abstract

atalyst tubes in steam reformers are made of high-alloy spun cast materials. One of the most important material properties for steam reformer tubes is the creep rupture

strength as it determines the tube lifetime. Schmidt + Clemens performed an intensive devel-opment program with the ultimate objective to ob-tain a material with superior stress rupture strength and creep resistance in order to improve the performance and productivity of steam refor-mers. The result is the alloy Centralloy® G 4852 Micro R. To achieve the best creep properties, stringent and specific quality control and quality assurance of the casting and melting processes have been im-plemented. The program involved different re-search and development projects in order to op-timize the key element composition, the nano-scale particle distribution, size, morphology and long-term stability with time.

Investigation results are presented as well as a number of advantages and operational benefits provided by the alloy Centralloy® G 4852 Mi-cro R.

Introduction

Steam reformers are used e.g. in the production of ammonia for fertilizers, hydrogen for oil refining, direct reduction of iron ore and syngas, a basis for the production of methanol, acetic acid and various other chemicals. In such steam reformers, an endothermic reaction takes place in vertical, catalyst-filled tubes that are directly fired. Because of the severity of the operating conditions, reformer tube assemblies are fabricated from centrifugally cast materials which provide excellent creep strength. The objective of a research program carried out in co-operation with the external research partner HDZ was to increase the stress rupture strength of the alloy Centralloy® G 4852 Micro (HP-MA).

C

65 AMMONIA TECHNICAL MANUAL2010

Materials for reformer tube assemblies

Damage mechanisms for reformer tubes

Catalyst tubes are subject to several stresses, e.g. - Hoop stresses generated by the pressure of the

gas within the tube, - Longitudinal stresses from the combination of

the axial component of the internal gas pres-sure and the tube weight and

- Thermally induced stresses due to the temper-ature gradient between the outer and the inner surfaces of the tube. These stresses are nor-mally tensile and act in the inner half of the wall thickness as longitudinal and circumfe-rential (hoop) stresses. They increase as wall thickness increases /1, 2/.

The combination of thermal stresses across the tube wall and internal pressure stresses causes creep damage that typically develops at the inner diameter (ID) or just below the ID surface. Also, creep damage occurs normally over a large part of the circumference and over a longer (axial) part of the tube. This process results in diameter increase and finally rupture would occur in a longitudinal direction. During the lifetime of a reformer tube, it will ex-perience a number of startups, shutdowns and op-erational transients; which generate very high through-wall stresses. Usually these are elastic, but at high rates of temperature change some plas-ticity may occur. Once the transient has passed, the stresses relax by creep towards steady state equilibrium. These high stresses lead to enhanced life consumption during this period of stress re-laxation and redistribution. The cumulative effect of several thermal cycles is very damaging and leads to accelerated creep cracking. Therefore, the tube life is mainly related to the number of startup and shutdown cycles it has seen.

Reformer tube life is also related to the tube wall thickness. Thicker tubes generate large thermal stresses during startup and shutdown and are thus significantly less resistant to thermal cycling than thinner tubes. The wall thickness of a catalyst tube can be opti-mized by the correct material choice, selecting one with high creep rupture strength.

Material choice for reformer tubes

Due to the severity of the operating conditions, re-former catalyst tubes are typically fabricated from the centrifugally cast HK (25Cr/20Ni), HP-Nb (25Cr/35Ni+Nb) or HP-MA (25Cr/35Ni+Nb+ MA) alloys. These spun cast materials contain around 0.4 weight % carbon and provide therefore excellent creep strength. The nominal composition of the materials is given in Table 1. The heat-resistant castings are general-ly known by the designation system adopted by the Alloy Casting Institute (ACI): The initial letter “H” indicates that the alloy is generally used for high-temperature service (metal temperature is in excess of 650 °C (1200°F). The second letter of the designation represents the nominal chromium-nickel type (“K” corresponds to 20Ni/25Cr and “P” to 35Ni/25Cr). Numerals following the letters indicate the maximum carbon content (e.g. HK-40). The abbreviation “MA” in Table 1 stands for “Micro- Alloying”. Because of the oxidation re-sistance at temperatures up to 1000 °C (1832°F), the chromium content is about 25%. This chro-mium content is the standard for all modern spun cast materials used as radiant tubes. The HK material does not use any carbide form-ing elements (except chromium) and develops primary M7C3-type carbides during casting, which are transformed into M23C6 during ageing. It is well established in the body of literature on the creep behavior of centrifugally cast alloys of HK- or HP-grade materials that micro-alloying with strong carbide formers like Ti, Nb or Zr can

66AMMONIA TECHNICAL MANUAL 2010

substantially influence the high temperature strength of the material beneficially /3-6 /. The newer HP-Nb alloys precipitate in addition to the mainly chromium containing M7C3- and M23C6 carbides also primary and secondary nio-bium carbides (NbC). In the early 1980´s, “micro”-alloys (MA) were developed and introduced. The most successful al-loy is the micro alloy based on the HP-composition (HP-MA). The micro alloy based on HK-40 has also been developed many years ago, but it has not become as popular. The HP-MA al-loy has proven to be stronger and the HK mate-rials in general tend to form the brittle σ- phase below 900 °C (1652°F). A sound overview with reliable experimental re-sults which still represent state of the art in this class of alloys is given in /4/. The important ef-fects of micro-alloying are (i) the primary precipi-tation of MC -carbides like NbC as part of the in-terdendritic carbide network additionally to M7C3 and (ii) the development of a secondary precipita-tion pattern inside the austenitic matrix with mono-carbides or carbo-nitrides (Nb, Ti, Zr)(C, N) which could be considerably smaller than sec-ondary M23C6 carbides. For technical applications it is important to tailor the particular additions of micro-alloying metals (V, Ti, Nb, Zr) and non-metallic elements (C, N) to obtain best and durable high temperature strength. Additionally, since an optimal precipita-tion pattern is the key requirement, it is desirable to fine tune the composition and the nature of sec-ondary carbides in such a way that they conserve their size and distribution over long periods of in-dustrial operation without significant ripening. In the field of High-strength low-alloy (HSLA)-steels the technique of dispersion particle streng-thening also has reached a matured technological level /7-9/ and at least concerning the basic metal-

lurgical background of precipitation strengthening there seem to exist a lot of interesting parallels. It was the aim of the present study to work out an improved concept of micro-alloying and of the corresponding processing of centrifugally cast tubes to obtain a superior high temperature ma-terial. The main results from this research program are highlighted below, together with its impact on fu-ture steam reformers.

Materials and experimental procedures

The chemical composition of the alloys investi-gated is given in Table 2. Samples for creep tests and for structural investigation are taken out of centrifugally cast tubes from the regular produc-tion. Alloys 277 and 278 are two variants of a HP-MA type of alloy (Centralloy® 4852 Micro) with different additions of micro-alloying elements Ti and Nb. The samples were selected from a series of batches in which the alloying fractions of Si, Ti, Nb and Zr were varied systematically. Sam-ples 277 and 404 represent batches with optimized compositions. Sample 278 was selected as exam-ple with an unfavorable Ti/(Ti+Nb)-ratio.

Sample Characterization

Samples were characterized at longitudinal metal-lographic sections. The micro-structure of samples in cast as well as in aged condition were analyzed by optical microscopy, and by field emission scanning electron microscopy with backscattered electron imaging BSE (FE-SEM, JEOL JSM 6330 F); for phase identification an Energy Dispersive X-ray analyzer was used (EDX, EDAX). The microscopy and microanalysis of the fine secondary precipitates was carried out in a high resolution analytical Scanning Transmission Elec-tron Microscope (STEM) (FEI TECHNAI F20 energy-filtering STEM operated at 200 kV). The

67 AMMONIA TECHNICAL MANUAL2010

instrument was equipped with a Gatan Imaging Filter (GIF) 2000, an EDAX-EDX spectrometer and a High-Angle Annular Dark Field detector (HAADF). EDX stands for Energy Dispersive X-ray spectroscopy. For the Transmission Electron Microscopy (TEM) investigation of the secondary precipita-tion after aging the extraction-replica technique was applied according to /9/. Before coating with a thin carbon film the polished samples were etched with austenite-pickle (100 ml distilled wa-ter, 100 ml HCl, 10 ml HNO3 0.3 ml Vogels Sparbeize).

Results

The as-cast tube material of the alloys generally exhibits a dendritic, directionally solidified ma-cro-structure (Figure 1) with typical precipitation of carbides, partly eutectic, on dendritic- and on grain boundaries (Figure 2a). These primary pre-cipitates consist of a mixture of chromium car-bides and monocarbides of the type NbC, this is due to the niobium additions. Depending on the amount of titanium addition to the alloy the primary MC-carbides may contain a small fraction of titanium as a substitute of Nb. Detailed inspection revealed no further precipi-tates inside the austenitic matrix in the as-cast condition. Under creep conditions at 1000 °C (1832°F) with 25 MPa (3626 psi) the eutectic primary carbides partly transformed from the lamellar into more globular shaped particles. This semi-spheroidization is more pronounced for the MC-carbides (Figure 2b). The Cr-carbides on the other hand, have been transformed from M7C3 to M23C6. Under creep conditions at temperatures 900 °C (1652°F) and 1000 °C (1832°F) the micro-structure of the alloys changes as follows (Figures 3, 4). Besides the transformations of the primary

carbides the austenitic matrix precipitates two dif-ferent types of secondary carbides. The first type consists of fine particles in the surroundings of primary carbides, which are secondary Cr-carbides of the type M23C6 precipitated from a su-persaturated, rapidly solidified matrix, see Figure 3. In the vicinity of the primary Cr-carbides their density is high and decreases towards the interior of the dendritic cells. This inhomogeneous distri-bution is caused by marked segregation of car-bide-forming elements in the interdendritic re-gions during solidification. It is this segregation zone where the secondary M23C6-carbides emerge preferably. Figure 4 shows in an extraction replica a typical seam of these particles. The secondary M23C6 partly have rectangular shape and seem to be semi-coherent or coherent with the face cen-tered cubic (fcc) structure of the matrix /5/. In sample 278 the secondary Cr-carbides are less frequent, especially around the primary MC-carbides. This can partly be explained by the higher addition of Nb in this alloy leading to more primary precipitation of NbC. The second type of precipitates could only be de-tected on a much finer scale by TEM. Figure 4 gives an overview image of a typical transition re-gion from the interdendritic carbides into the cen-ter of a cell. Very small particles could be de-tected which are distributed over the whole matrix and appear also between the secondary as well as between the primary Cr- carbides. The density of the particles is high and they seem to be nucleated at dislocations, subgrain boundaries or even at stacking faults. These particles could further be characterized by Electron Spectroscopic Imaging (ESI) (energy fil-tered TEM) and EDX-analysis, and were found to be Nb-Ti-compounds. Electron diffraction identi-fied most of them as having a fcc crystal structure according to NbC. A detailed series of EDX- and ESI analyses revealed that these nano-sized preci-pitates do not have a constant but rather a varying

68AMMONIA TECHNICAL MANUAL 2010

Nb:Ti – ratio. By the ESI-analyses shown in Fig-ure 5 and Figure 6 it could be proved that the na-no-sized precipitates are heterogeneous: Many of the small particles have separate Nb-rich and Ti-rich parts, respectively; in that the Ti-rich parts contain nitrogen. It can be assumed that the Ti-rich particles are Ti-Nb-carbonitrides which acted as nuclei for the precipitation of the Nb-rich par-ticles. The latter seem to have small fractions of Ti and are (Nb, Ti)C-carbides. In the center of Figure 6 two Nb-rich carbides are nucleated at a large Ti- and N-rich particle. Details on the exis-tence of ternary carbide phases in the system (Nb, Ti)(C, N) are discussed in /10/. Turning to the mechanical properties, Table 3 summarizes data obtained from creep tests and re-sults of a particle-size analysis of the nano-sized secondary mono-carbides for the alloys investi-gated. It turns out that the performance of the al-loys in the creep test directly depends on the atomic ratio Ti/(Ti+Nb). For the alloys with the higher ratio the density of nano-sized mono-carbides in the matrix is very high. Sample 277 has a particle density of 10 µm-² (areal particle density determined from TEM-micrographs of ex-traction replicas), in case of sample 278, with a lower Ti fraction the density drops to 4 µm-². As a result, sample 278 reveals an almost doubled min-imum creep rate at 1000 °C (1832°F) leading to a lower time to rupture. Thus, the measured particle densities of the secondary mono-carbides directly correlates with high-temperature strength of the alloy. Alloy 404, with a similar composition as al-loy 277 was creep tested at 900 °C (1652°F). In this case the particle density of the secondary mono-carbides increased by 30% compared to 277, which indicates a higher density of nuclea-tion sites at the lower temperature. For all samples the average diameter of these particles is in the range between 20 and 35 nm (measuring uncer-tainty considered), thus it seems to be independent of the differences in creep temperature or the du-ration of tests.

According to the TEM results from the extraction replicas MC particles with diameters less than 10 nm could be detected.

Benefits available by installation of Centralloy® G 4852 Micro R

With these new results at hand, the alloy Central-loy® G 4852 Micro R has been invented, combin-ing all the beneficial effects described above. In the typical steam reformer conditions, the stress rupture strength determines the lifetime of the cat-alyst tubes, or, on the other hand, determines the Minimum Sound Wall thickness (MSW) required to design a tube. This MSW can be calculated by simple equations such as given e.g. in API RP-530 (Calculation of Heater Tube Thickness in Petroleum Refineries): MSW=P*OD/ [(2Sa)+P]. It takes into considera-tion the internal pressure stresses (design pressure P), the tube diameter (OD) and the allowable stress (Sa) which is defined as minimum 100,000 hours creep rupture stress – but it does not take in-to account the thermal stresses caused by start/stop cycles. This minimum 100,000 hours strength corresponds to the lower limit of the scat-ter band (95% confidence interval). The characteristic Larson Miller curve (parametric stress rupture strength) for the alloy Centralloy®

G 4852 Micro and the recently improved material Centralloy® G 4852 Micro R is shown in Figure 7. This improved alloy Centralloy® G 4852 Micro R now offers our customers a range of benefits or combinations of benefits from which to choose. These include:

Increased lifetime

By retaining the existing tube design, (dimension-al parameters) and operating conditions, this alloy substantially increases the lifetime of the tubes.

69 AMMONIA TECHNICAL MANUAL2010

Process improvements

By maintaining existing reformer tube dimensions (except for using the higher strength of this alloy to reduce the MSW, and the subsequent increase in tube ID), allows process gains through in-creased catalyst volume and improved heat trans-fer. Results of exemplary calculations are given in Table 4. The reduced MSW results additionally in an increased resistance to thermal shock. Alternatively, if tube dimensions remain un-changed and the design remains based on the standard 100,000 hours, the higher creep strength of Centralloy® G 4852 Micro R allows an increase of more than 20 °C (36°F) in operating tempera-ture, which will result in significant capacity gains. Commercial advantage: The thinner wall of tubes in Centralloy® G 4852 Micro R means less weight and less cost, especially when Nickel in particular is trading at high prices, as seen not so long ago. Starting some years ago S+C carried out an inten-sive program of testing Centralloy G 4852 Micro R in a number of operating plants and the alloy proved completely successful in meeting all ex-pectations. Since 2007 it has been commercially available and installed worldwide in all types of reformer furnaces e.g. hydrogen, ammonia and methanol, and to many different furnace designs, e.g. Lurgi, Uhde, Technip/KTI ,Linde, Kellogg and Foster Wheeler. We are pleased to confirm that in all cases, results to date meet or exceed initial expectations.

Conclusions

The creep resistance of the alloy Centralloy® G 4852 Micro could be improved substantially by tuning the additions of micro-alloying elements as

well as non-metallic elements and solid-solution hardeners in a synergistic way.

It could be worked out that an addition of Ti acts threefold, it stimulates the nucleation of secondary precipitation of Nb-Ti-carbides or –carbonitrides, it stabilizes their size and dispersion, and Ti is implemented into the mono-carbides as a substi-tute of Nb.

The enhanced nucleation seems to be correlated with the availability of carbon and nitrogen in the matrix. The nuclei could be identified as small Ti-Nb-carbonitrides.

Thus, besides the optimal Ti:Nb – ratio also a suitable amount of nitrogen should be available in the matrix of the as-cast material in fact in super-saturated solid solution.

Regarding to the micro-alloying additions Nb and Ti the optimized alloy has best creep performance with an atomic ratio Ti/(Ti+Nb) of 0.16 producing the following beneficial effects:

(i) Spheroidization of primary MC-carbides from lamellar to more globular shape,

(ii) Strong increase of the number of secondary nano-sized mono-carbides or carbonitrides of type (Nb, Ti)C or (Nb, Ti)(C, N) leading to an im-provement in creep resistance (decrease of mini-mum creep rate by almost 50%),

(iii) Improvement of stability and dispersion of these nano-sized particles: with Ti addition the average size after 2000 h at 1000 °C (1832°F) was smaller by 50% compared to an alloy without Ti. In order to achieve the creep properties in the reg-ular production, stringent and specific quality con-trol and quality assurance of the casting and melt-ing processes have been implemented. This recently improved material Centralloy® G 4852 Micro R now offers our customers a range of benefits like improved lifetime and/or process improvements as e.g. improved heat transfer or significant capacity gains due to an increased

70AMMONIA TECHNICAL MANUAL 2010

catalyst volume or an increased operating temper-ature.

References

1. R. Gommans, D. Jakobi, J. L. Jimenez; State-of-the-art of materials and inspection strategies for reformer tubes and outlet components, 46th An-nual Safety in Ammonia Plants and Related Facil-ities Symposium, Montreal, Canada, 2001 2. T. Kawai, M. B Zaghloul; Design Features, Material Selection and Performance of Steam Re-forming Furnace, International Plant Engineering Conference (INPEC), Bombay, November 1984 3. L. T. Shinoda, M. B. Zaghloul, Y. Kondo; The Effect of Single and Combined Addition of Ti and Nb on the Structure and Strength of the Centrifu-gally Cast HK-40 Steel; Transactions ISIJ, 18 (1978) 139 4. Hou Wen-Tai, R. W. K. Honeycombe; Struc-ture of centrifugally cast austenitic stainless steels: Part 2 Effects of Nb, Ti and Zr; Materials Science and Technology, 1 (1985) 390 5. L. H. de Almeida, A. F. Ribeiro, I. Le May: Microstructural characterization of modified 25Cr – 35Ni centrifugally cast steel furnace tubes; Mat. Characterization 49 (2003) 219

6. F. C. Nunes, L. H. de Almeida, J. Dille, J.-L. Delplanke, I. Le May: Microstructural changes caused by yttrium addition to NbTi-modofied cen-trifugally cast HP-type stainless steel; Mat. Cha-racterization 58 (2007) 132 7. A. J. DeArdo, Mingjian Hua, C. I. Garcia: Ba-sic metallurgy of modern niobium steels; Interna-tional Symposium on Niobium Microalloyed Sheet Steel for Automotive Application, Edts. S. Hashimoto, S. Jansto, H. Mohrbacher, F. Sicilia-no, TMS (The Minerals, Metals & Materials So-ciety), 2006, page 499 – 549 8. C. Klinkenberg, K. Hulka, W. Bleck: Niobium Carbide Precipitation in Microalloyed Steel; steel research int. 75 (2004) No. 11 1 – 9 9. M. Béres, T. E. Weirich, K. Hulka, J. Mayer: TEM investigations of fine niobium precipitates in HSLA steel; steel research int. 75 (2004) No. 11, 10 -15 10. K. Inoue, N. Ishikawa, I. Ohnuma, H. Ohtani, K. Ishida: Calculation of phase equilibria between austenite and (Nb, Ti,V)(C,N) in microalloyed steel; ISIJ International 41 (2001) 175

71 AMMONIA TECHNICAL MANUAL2010

Table 1 Nominal chemical composition of centrifugally cast materials for reformer catalyst tubes Material type (ACI designation) Centralloy® G C Cr Ni CFE* HK 4848 0.45 25 20 - HK-MA 4848 Micro 0.45 25 20 Nb, Ti, Zr HP 4852 0.45 25 35 Nb HP-MA 4852 Micro 0.45 25 35 Nb, Ti, Zr HP-MA 4852 Micro R 0.45 25 35 Nb, Ti, Zr

CFE*= carbide forming elements Table 2 Chemical composition of the alloys (wt.-%) Alloy C Mn Cr Ni Fe CFE Ti/

(Ti+Nb)* 277 0.42 0.98 25.3 33.2 Bal. Nb, Ti, Zr 0.16 278 0.43 0.92 24.8 33.6 Bal. Nb, Ti, Zr 0.04 404 0.39 0.83 25.2 35.7 Bal. Nb, Ti, Zr 0.16

*: atomic ratio

Figure 1 Sample 277, optical micrograph of the microstructure showing a typical dendritic solidification struc-

ture.

72AMMONIA TECHNICAL MANUAL 2010

(a) (b) Figure 2

Sample 277, SEM-BSE images of the microstructure with primary carbide network: (a) as-cast condition, (b) after creep test, 1000 °C (1832°F), 25 MPa (3626 psi).

(a) (b) Figure 3

SEM-BSE images of the microstructure after creep testing, 900 °C (1652°F), 40 MPa (5802 psi): (a) sample 277, (b) sample 278.

(Nb, Ti)C

M23C6

M23C6

secondary M23C6 secondary M23C6

(Nb, Ti)C

M7C3 M23C6

(Nb, Ti)C

(Nb, Ti)C

73 AMMONIA TECHNICAL MANUAL2010

secondary

primary

carbide

secondary M(C N)

500 nm

Figure 4 TEM micrograph of typical precipitates in alloy 404 after a creep test at 900 °C (1652°F):

Carbon extraction replica. Areas of primary carbides are indicated grey.

Table 3 Results of creep tests and properties

Alloy Temperature [°C] (°F)

Applied stress

[MPa] (psi)

Creep-rupture time [h]

Minimum creep rate

[% h-1]

Density of nano-sized MC/µm²*

Average diam. of nano-sized

MC [nm]

Smallest par-ticle size de-tected [nm]

277 1000 (1832) 25 (3626) 1925 1.6x10-4 10 27 9 278 1000 (1832) 25 (3626) 1458 2.9x10-4 4 28 9 404 900 (1652) 40 (5802) 10327 7x10-5 13 27 3

*: particles/µm² determined from extraction replica TEM-micrographs

74AMMONIA TECHNICAL MANUAL 2010

Bright Field TEM micrograph titanium niobium

Figure 5 Micro-chemical characteriza-tion of secondary precipitates by Electron Spectroscopic Imaging (ESI): Large particles: Cr-carbides; Small particles: NbC nuc-leated at Ti(C,N).

chromium nitrogen

Figure 6 Micro-chemical characteriza-tion of secondary precipitates by Electron Spectroscopic Imaging (ESI): Small particles: (Nb, Ti)C nucleated at (Ti, Nb)(C, N).

Bright Field TEM micrograph niobium

titanium nitrogen

100 nm

Ti(C, N)

Ti(C, N)

NbC

M23C6

Ti(C, N)

Ti(C, N)

NbC NbC

75 AMMONIA TECHNICAL MANUAL2010

1

10

100

800 850 900 950 1000 1050

Temperature, °C (°F)

Initi

al S

tres

s, M

Pa (p

si)

4848 (HK-40)4852 (HP-Nb)4852 Micro (HP-MA)4852 Micro R

(1472) (1562) (1652) (1742) (1832) (1922)

Minimum 100,000 h Creep Rupture Stress (14504)

(1450)

(145)

Figure 7

Parametric stress rupture strength of alloys Centralloy® 4852 Micro and Centralloy® 4852 Micro R. Table 4: Influence of the creep rupture strength on the minimum sound wall thickness (MSW), catalyst volume and tube weight (design parameters: P = 4.0 MPa (580.15 psi); T = 925 °C (1697 °F); OD = 114.3 mm (4 in)

Centralloy® G min. stress to rupture

at 100,000 h [MPa] (psi)

MSW [mm] (in)

ID [mm] (in)

catalyst volume

tube weight [kg/m] (Lb/in)

4852 19.0 (2756) 10.9 (0.429) 92.5 (3.64) reference 28.3 (1.58) 4852 Micro 22.0 (3191) 9.5 (0.374) 95.3 (3.75) + 6.1% 25.0 (1.40) 4852 Micro R 25.5 (3689) 8.3 (0.327) 97.7 (3.85) +11.6% 21.8 (1.22)

4852

4852

Mic

ro

4852

Mic

ro R

15

18

21

24

27

Min

. Stre

ss, M

Pa

(psi

) (3916)

(3481)

(3046)

(2611)

(2176)

4852

4852

Mic

ro

4852

Mic

ro R

7.0

8.0

9.0

10.0

11.0

MSW

, mm

(in)

(0.275)

(0.315)

(0.354)

(0.394)

(0.433)

4852

(ref

eren

ce)

4852

Mic

ro

4852

Mic

ro R

0%

2%

4%

6%

8%

10%

12%

Cat

alys

t Vol

ume

Incr

ease

min. 100,000 h stress rupture strength minimum sound wall thickness catalyst volume increase

76AMMONIA TECHNICAL MANUAL 2010