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Philli p Dowson is General Manager,

Materi als Engineering, with Elli ott Company,

in Jeannette, Pennsylvania. He has 35

years of experi ence i n the tur bomachinery

industry. M r. D owson i s responsible for the

metallurgi cal and welding engineeri ng for

the various Elliott product lines within the

company. He is the author/coauthor of a

number of technical arti cles, related to topics

such as abradable seals, high temperature

corrosion, fr acture mechanics, and weldi ng/brazing of impellers.

Mr. Dowson graduated from Newcastle Polytechnic in Metall urgy and did his postgraduate work (M.S. degree) in Welding Engineering. He is a member of ASM, NACE, ASTM, and TWI.

Derrick Bauer i s a Materi als Engineer with

Elliott Company, in Jeannette, Pennsylvania .He joined Elliott Company in 2002, and has been involved wi th materials related R&D projects, fail ure analysis, production and aftermarket support, and remaining life assessments.

Mr. Bauer received his B.S. degree (2002)

from the University of Pittsburgh and is

currently working toward his M.S. degrees

from the same institution.

Scot Laney is a Materials Engineer

with Ell iott Company, in Jeannette,Pennsylvania. H e j oined Ell iott Company in 2007, and has been involved with materi als related R&D projects, failure analysis, and aftermarket support. He also has experience in the areas of high temperature oxidation/corrosion and protective coati ngs.

Dr. Laney received his B.S. degree (2001), M.S. degree (2004), and Ph.D. degree (2007) from the University of Pittsburgh in M ateri als Science and Engineeri ng. He is al so a member of ASM.

ABSTRACTIn today’smarketplacethe selection of materials for the various

componentsfor centrifugal compressorsandsteamturbinesisverycompetitiveandanimportantfactor in theoverall cost anddeliveryof the product. This paper reviews thematerial selection for majorcomponents for compressors and steam turbines such as shafts,impellers, blading, bolting, seals, etc. This paper is not intended tocover reciprocating, screw compressors, and high temperaturehotgas expanders. Various material properties are discussed for themanufacture and service exposure of major components such asimpellers, shafts, bolts, blades, and casings. Due to the variousaggressive corrosive, fouling, and erosive environments in whichboth compressors and steam turbines are operated in, coatingsmust frequently be used to prolong the life of the machine. Theapplication of various coating systems will be reviewed and how

effective the coating system is with respect to withstanding theparticular environment. Also discussedaretherepairs for longleadtimecomponentssuchasrotorsandtheapplicationof designprovenrepair procedures utilizing adesign for fitnessfor serviceapproach.

INTRODUCTION

Over thepast15yearsagreat deal of progresshasbeenmadewithrespect to the application of materials and related processes appliedto various components for centrifugal compressors and industrialsteam turbines. This paper will review how materials are specifiedfor thevariouscomponentsandwhichmaterial properties/factorsonemust consider for the application. Both rotating and nonrotatingcomponents of centrifugal compressors and steam turbines will bereviewed for material selection.

Due to the very competitive market, material selection has

moved beyond simply finding the material with the most idealproperties. Material cost and delivery can be one of the mostimportant factors in the overall cost and delivery of the product,andthereforehavebecomemajor drivers when selectingmaterials.Most original equipment manufacturers (OEMs) are continuouslyreviewing new ways whether by material changes or processes toreduce costs or delivery in order to remain competitive. Thisbecomeseven more important with the prevalence of outsourcingand inventory concepts such as “just in time.” OEMs are nowcompeting not only for sales, but also as customers in places likecasting andforging shops. At the same time, the environments thecomponentsareexposedtoareincreasingin severity, leadingtothe

189

SELECTIONOF MATERIALSAND MATERIALRELATED PROCESSES FOR CENTRIFUGAL COMPRESSORS

AND STEAM TURBINES IN THE OIL AND PETROCHEMICAL INDUSTRY

by

Phillip Dowson

General Manager MaterialsEngineering

Derrick BauerMaterialsEngineer

and

Scot Laney

MaterialsEngineer

ElliottCompany

J eannette, Pennsylvania



need for more specialized materials. The use of these typicallymore costly materials may be avoided by selection of variouscoatings that are resistant to the aggressive environment underconsideration. As statedpreviously, thepaper will reviewmaterialsselectionsfor:

•Centrifugal compressors—bothrotatingandstationary components.

•Industrial steamturbines—rotating and stationary components.

•Applicationof coatings for protection against corrosion, erosion,andfoulingenvironments.

CENTRIFUGAL COMPRESSORS

The choice of materials for rotating and stationary componentsof centrifugal compressorsrequiresselectionbasedonanumber of design and environmental factors. The design factors that requireconsideration are the properties of the materials together with theoperating requirements of the unit. For example, operating amachine intermittently or with variations is usually considered amuch more severe service relative to constant, long-durationservice, and must be designed accordingly. The properties of materials that can be utilized for usein the design areas stated byCameron and Danowski (1973) and listed in Table 1. Long time,high temperature properties such as creep or creep rupture arerarely encountered in centrifugal compressors and, therefore, arenot considered in this selection process. However, in cases where

the temperature of operation is determined to be signif icantlydifferent from room temperature, temperature is an importantfactor, since the following properties are temperature dependent:strength of material, corrosion rate, coefficient of thermalexpansion, and fracturetoughness.

Table 1. List of M ateri al Properti es Used i n D esign of Centrifugal Compressors.

From an OEM perspective, receiving quality information withrespect totheanticipatedoperatingparametersisvital toproducingacompressorthat meets thecustomer’sexpectations. API Standard617(2002)providessomeguidelinesastowhatinformation shouldbe shared by both the OEM and the purchaser. For instance,paragraph 2.2.1.3 requires the purchaser to specify any corrosiveagentsthat may bepresent. Asanexampleof why this is important,thereis asubstantial differencebetweenthematerials thatareusedin an air compressor when compared to an application that maycontain wet chlorine.

NACE MR0175 (2003) and MR0103 (2005) define whichferrousandnonferrous alloys can beused in wet hydrogen sulf ideservicetoresistsulf idestresscorrosioncracking. Thealloysallowedby thesespecificationsmust also beheat treated in accordancewith

the specifications and meet a maximum hardness limit. NACEMR0175 (2003) is labeled “Metals for Sulf ide Stress CorrosionCracking and Stress Corrosion Cracking Resistance in SourOilfield Environments,” which applies to a component ormachinery used for petroleumproduction, drilling, flowlines, andfield processingfacilities exposed towet hydrogensulf ideservice.NACE MR0103 (2005) is titled “Materials Resistant to SulfideStress Corrosion Cracking in Petroleum Refining Environments,”andis thus morespecific to compressor components. Compressorswith wet hydrogen sulfide present in the process gas are usuallyrequiredtomeet oneorbothof theseNACE specificationstoavoidstresscorrosioncrackingduring service.

Rotating Components

Theheartof acentrifugal compressor istherotor, whichconsistsof a series of impellers and a shaft. Theimpellers are designed toacceleratetheprocessgas, which causes it tobecompressed in theproceeding diaphragm, while the shaft provides the support androtation to the impellers.

Impellers

Given the importance of theimpellers, a great deal of attentionis given to their manufacture. Generally, centrifugal compressor

impellers are fully shrouded, consisting of a solid hub and coverseparated by radial equally spaced blades. The attachment of thebladestotheimpeller hubandcover canbeeither donebywelding,integral cast, brazing, riveting, electrodischargemachining, integrallymachinedtothehuband/orcover, oracombination of thesemethods.A fully shrouded impeller is shown in Figure 1. Today, for mostimpellers, the blades are integrally machined to the impeller hub orcover andthen welded to thenonmachinedbladehub or cover.

Figur e 1. Photograph of a Full y Shrouded Centri fugal Compressor Impeller.

Throughout the compressor industry selection of impellermaterials can vary depending upon service operating conditions.Sinceoperatingconditionsgovern thematerial selection requirements,anumber of mechanical and chemical properties suchasyield andtensile stresses, low temperature properties, corrosion resistance,hydrogen sulphide resistance, weldability, and machinability must

be considered. For over 50 years, OEMs have utilized impellermaterials such as AISI 4330/4340, 4130/4140, AISI 410/17-4PHand13Cr4Ni and nickel (Ni) basealloy.

Impellers for centrifugal compressors are typically made fromlow alloy steels or stainless steels, andbothNACE MR0175 (2003)and MR0103 (2005) have identical requirements for commonimpeller materials. I mpellers made from AISI 4140 or 4130 steelhave to meet a maximumhardness requirement of 22 HRC in thebase metal, weld metal, and heat-affected zone (HAZ). In order toachieve the necessary hardness requirement, AISI 4140 impellershave to be austenitized, quenched, and tempered after welding toeliminatetheHAZ. UNS G43200andmodif iedversionsof G43200with higher carboncontentsareacceptablefor compressor impellersby both NACE MR0175 (2003) and MR0103 (2005), provided thatthey are heat treated per the NACE standards. TheseUNS G43200have a maximum yield strength requirement of 90 ksi, but there isno hardness limitation given per the NACE standards. AISI 410stainless steel meets both applicable NACE standards if it isaustenitized and double tempered with the second temper beingperformed at alower temperaturethan thefirst temper. BothNACEMR0103 (2005) and MR0175 (2003) allow the use of low carbonstainless steel CA6NM (13%Cr-4%Ni) for impellers at a maximumhardness of 23 HRC after the material has been austenitized anddouble tempered at the temperature ranges defined in the NACEspecif ications. Precipitation hardenable stainless steels UNS17400 (17-4PH) and UNS15500 (15-5PH) also meet both NACEspecifications at a maximumhardness of 33 HRC after it has beenthrougha solution annealingfollowed by a double agingtreatment.

PROCEEDINGSOF THE THIRTY-SEVENTH TURBOMACHINERY SYMPOSIUM •2008190



Since the maximum hardness requirement is 33 HRC for thesesteels, they are often used when impellers with higher yieldstrengths arenecessary in hydrogen sulfideserviceenvironments.

Although other manufacturing techniques havebeen utilizedin thepast20years, most impellersareweldedusing variousweld processes.Other manufacturingtechniques that havebeen utilized are:

•Castimpellers.

•Electrodischarge machining to shape the gas passage of impellers.

•Riveted impellers.

•Brazing impellers.

•Electronbeamwelding.

•Combination of electronbeamweldingandbrazing.

•Onepiece machineimpellers.

Early on, shieldedmetal arc wastheweldingprocessutilizedfor joiningthesectionsof impellers, whether it wouldbemilledbladestocover/hubtotheunmilledsectionor joiningpreformedbladestoamachined hub andcover.This weldingprocessis still used today.In the 1980s several manufacturers started to utilize automationwith other welding processes such as gas tungsten arc welding(GTAW), gas metal arc welding (GMAW), and submerged arc

welding (SAW). Generally, theseprocesses wereapplied to simpletwo-dimensional (2D) configuration impellers with 2D curvedblades. Special torches were designed for both open bladed andclosed configuration welding.

Due to technological advances in five axis milling, two-piecethree-dimensional (3D) geometric compressor impellerswithvarioustypes of three-dimensional curved blades were being designed andmanufactured. Theseimpellersweredesignedtoincreaseperformanceof the wheel. Since the impellers are two-piece designed, that is,either the blades are milled to the hub or cover, material selectionbecomesmoreimportant for consideration of manufacturability.

Brazing of impellers has been used since the 1950s but in therecentyearsits popularity hasincreased. Early brazedimpellersweredone in a dry hydrogen atmosphere. More recently, however,impellers are done in a vacuum furnace, which produces more

acceptable results andconsistency. Theinspectionof thebrazejointshas also improved over the years with the application of C-scanimmersionultrasonic testing.Theauthor’scompanyutilizesaspecialcalibration block to represent the impeller braze joint. This specialblockwasutilizedfortheapplicationof C-scanultrasonic immersionequipment. TheFigure2 shownis aprintoutof an acceptable C-scanultrasonic immersion test. Brazed impellers are also fluorescentpenetrant inspected to validatethesoundnessof thebrazejoint.

Figure 2. Immersion C-Scan U ltrasonic Results from a Brazed Impell er Joint.

Materials such as low alloy steels have limited hardenability.Hardenability also becomes an issue as the diameter of the

impeller increases.This is illustrated in Figure3, which showsthehardenability curvefor varioussectiondiameters of AISI 4140. Tounderstandwhy material mechanical properties suchasAISI 4140,4340, and 4330 diminish from the surface to the center of thematerial, one must review the continuous cooling transformationdiagram (CCT) for that particular alloy and an elementaryunderstanding of heat transfer. Figure 4 shows the CCT for AISI4140(Atkins, 1980). Dependingontherateof cooling, whether byair, oil, polymer, or water, for various thicknesses of material,certain microstructure phases such as ferrite, pearlite, bainite, and

martensite are formed. Each one of thesephases has an effect onthe strength, ductility, and toughness of the material. Table 2(Cameron, 1989) shows mechanical properties of impellermaterials. Consequently, special consideration must be appliedwhen selectingtheselow alloy steels in heavy forgingsfromwhichthe blades are milled. For example, during austenitizingtreatmentto a fabricated impeller followed by a rapid quench, specialattentionmust bemadeto changesinsectionthickness, suchasthatwhichoccurswherethebladesareweldedtoeither thehuborcoverin order to minimize the risk of quench cracking at the blade tohub/cover location.Insomecases, quenchingby forcedair-coolingorsalt bathmay berequiredto avoidthequenchcrackingphenomenon.Other high alloy materials such as AISI 410, 17-4PH, and 13Cr4Nihavegoodhardenability withlessdrastic coolingrates; requiringonlyair or forced air coolingtoachievethedesired mechanical properties.Figures 5 (A tkins, 1980) and 6 (Arlt, et al., 1988) show the CCTdiagrams for AISI 410 and13Cr4Ni, respectively.

Figure 3. Hardenability Curves for Standard and Controlled Yield

Strength AISI 4140.

Figure 4. CCT Curves for AISI 4140.

SELECTION OF MATERIALSAND MATERIALRELATED PROCESSES FOR CENTRIFUGAL COMPRESSORS

AND STEAM TURBINESIN THE OIL AND PETROCHEMICAL INDUSTRY

191



Table 2. Mechanical Properties of Impeller Materials.

Figure 5. CCT Di agram for AISI 410.

Figure 6. CCT Diagram for 13Cr4Ni.

The material 13Cr4Ni (UNS S42400) has been utilized formanufacturing impellers to meet a wide range of requirementssuch as high strength, controlled hardness H2S, low temperature,and good corrosion resistance. For low temperature applications

using 13Cr4Ni, controlling the volume fraction of austenite in thestructure increases the material toughness. This increase intoughnessalso can lead toareductionin hardness. Dowson (2002)also substantiated the application of a two-stage temperingtreatment to obtain the maximum requirement of HRC 23maximum. For welding, matchingweldingconsumablesarealwaysappliedto higher alloysteels.Theauthor’scompanyhasformulateda patented chemistry for the welding consumables of 13Cr4Ni tobe applied for impellers. This patent chemistry enabled HRC 23maximumhardness to beachieved consistently for H2S controlledhardnessenvironments. Proceduresweredevelopedusing13Cr4Nifor the various operating conditions utilizing the various welding

processes and brazing application. These procedures weredeveloped for welding and brazing applications requiring highstrength, qualification to meet NACE MR0175-200 (2003)(metallic sulf ide stress cracking resistant materials), and lowtemperature(downto166F (110C) applications. TheFigures7and8 showthepropertiesthat canbeachievedusingthe13Cr4Nimaterial (Dowson, 2002).

Figur e 7. Plot of Strength and Har dnessVersusTempering Temperature

(4 Hour H old) for 13Cr4Ni.

Figur e 8. Plot of Charpy Impact Energy VersusTemperi ng Temperature (4 Hour H old) for 13Cr4Ni.

TheprecipitationgradessuchasArmco 17-4PH,Armco 15-5PH,and Carpenter Custom 450 obtain their mechanical properties bysolution treatment and precipitation treatments. By varying theprecipitation treatments, onecanobtainarangeof tensilestrengths.In comparing the A rmco 17-4PH with A rmco 15-5, whosechemistriesoverlap,17-4PH is moreproneto formdeltaferritethan15-5PH, which may result in a loss of ductility/toughness. Theprecipitation treatments for these grades are usually in the range of 1050F to 1175F. For applicationinhydrogenandhydrogensulf ideenvironments,17-4 or15-5PH material is givenadouble precipitationtreatment at approximately 1150F to obtain the necessarymaximum hardness of HRC 33 maximum for H2S environments

and120 ksi maximumyield strengthfor hydrogen environments.

Special Impell er M ateri als

There are other special materials that are utilized in certainspecialized applications. In moist halogen machines, such as achlorine machine, Monel® K500 has been successfully used forimpellers. Monel® K500 was also used in oxygen machines,because of its resistanceto sparking. Special welding procedureswere developed to obtain good quality fabricated Monel® K500impellers. Titanium and titanium alloys have been used for wetchlorineandin special cases wherelow density makes thematerialattractive. Precipitationhardenable nickel basealloyssuchas UNS




N0 7716 and UNS NO 7725 have been applied to aggressivecoal gasification applications with a great deal of success. Fortemperatures down to 320F, the 9%Ni steel has widely beenaccepted for impellers in compressors for boil off gas fromliquidmethaneduetoitshighfracturetoughnessattheselowtemperatures.Special processing of these grades is required to obtain therequired mechanical properties. For some applications, aluminumalloys have also been used successfully especially in connectionwith liquefiednatural gas projects.

Rotor Shafts

Generally, shafts are manufactured from either rolled bars orforgings. In the size range where both are acceptable by API617 (2002) (currently inches finished machined), the choicebetween rolled bar or a forging should be dependent on costand availability, as there is little difference in the end result, if manufactured to relevant American Society for Testing andMaterials (ASTM) procedures. Tolerances on rolled bar can bemore tightly controlled, which results in less machining requiredthan for forgings and havea slight advantage in terms of thermalstability. Conversely, depending on availability of the particularmaterial andsize, selecting rolled bar may require the purchaseof an entire length of bar or an entire heat, while forged shafts are“made to order.” Shafts with a finished diameter greater than 8inches are generally madefrom forgings.

For shaft materials thestandardAISI 4330and4340can beheattreated to give the required mechanical strength and toughnessvalues. Due to these materials having higher strength values, thelateral expansionparameter is utilizedforASME Boiler andPressureVessel Code Section VIII Division I (2007). The requirement is 15mils, which was proposed by Gross and Stout (1957). Thesematerials have been used successfully down to 150F with thetensile strength at 110 ksi minimum and yield strength of 90 ksiminimum. For temperatures lower than 150F, ASTM A470Class7 has been used with a great deal of success.

Shaftsthat areused in H2S environmentcannotbelimitedtotheNACE requirements due to the need for the higher strengthrequired at the drive end of the shaft. Consequently, the appliedstress in the main body of the shaft, which comes in contact withthe gas, is low (<10 ksi). This valueis below the threshold stressvalue for sulfide cracking as a function of hardness (Figure 9)(WarrenandBeckman1957).

Figure 9. Threshold Stress for Sulfi de Cracking as a Function of Hardness.

Theselectionof shaftmaterialsisalso important to prevent certaintypes of failures, in particular, wire wooling. Wire wooling is aninfrequent, but devastating, failure mechanism that occurs duringstartup(not necessarily theinitial startup) in thebearingandseal areashafts. Areas whereoil f ilmsarethin and loads arehigh, such as at

the thrust bearing, are most susceptible to this type of failure. Theprocess of wire wooling begins when a small particle of foreignmaterial enters the bearing or seal. Through a series of localizedtemperature increases, due to high coefficient of friction betweenparticle and rotor, material transfers from the rotor to the particle,and by hardening mechanisms, the particle becomes a hard, blackscab that is able to cut and spin material from the shaft. Shaftmaterial also continues to transfer to the scab as the cutting isoccurring, allowing the scab to grow andpropagatethe failure. Theresult is adeeply groovedshaft. Thematerial cut fromtheshaft often

has the appearance of wire wool. Figure 10 shows damage to the journal and thrust collar due to wire wooling and the correspondingblack scab onthejournal bearing. For amoredetailed description of the mechanism, the reader is directed to Fidler (1971).

Figure 10. PhotographsShowing Wire Wooling Fail ure of a Journal and Thrust Collar and the Corresponding Black Scab on the Journal Bearing.

For reasonsthat arenot clearly understood, shaft materials withhigher Cr contents are more susceptible to wire wooling. Thecritical Cr content proposed by Fidler (1971) is 1.8 percent.Because of this, shafts should bemadeof 4140or 4340whereverpossible. If a higher Cr content material must beused, stepsmustbetaken toreducetherisk of wirewooling. Alteringthesurfaceof the shaft or increasing clearances are ways that can be used toreduce the risk. One way to alter the surface of the shaft is toharden it by nitriding or hardfacing with hard chrome or carbidecoatings. Sleevingwith a wire wooling resistant material, such as4140, can also work. Increasing the clearances reduces risk by

increasing the size of particle able to initiate the failure. All of these have their pros and cons, which must be evaluated whendecidingwhich methodyieldsthe bestoverall results.

Stationary Parts

Thebulk of thecentrifugal compressor consistsof stationaryparts. Themost obviousis thecasing, whichis thepressurecontainingshellthat surrounds the rotor. Diaphragms are responsible for slowingdownthe gas after it is accelerated by the impellers, causingit to becompressed. Bolts are used to fasten the parts together. Split lineseals fit between the casing pieces to maintain the overall pressure.Abradable andrub tolerant seals maintain thepressurebetweeneachindividual stage. Pipingis used for gas path connections.

Casings

Nearly all multistage centrifugal compressor casings are madefromcarbonor a low alloy material that is either cast or fabricatedfrom castings, forgings, plate, or a combination of these.Occasionally higher alloy steels are required for aggressivecorrosionconditionsor toachievethedesiredtoughnessfor extremelow temperature (320F) operating conditions, such as those forboil off gas compressors in liquified natural gas service. For lowtemperature application down to 175F, low alloy steels such asshown in Table 3 are applied in propaneandethylene compressors.

Thetable must be used with some caution and understanding of theeffect chemical makeup will have on the ability to achieve thedesiredtoughness.



193



Table 3. Mi nimumTemperature for Low TemperatureAppli cation of Low Alloy Steels.

Compressor casings are typically manufactured from carbonsteel castings or carbon steel plate that is formed and welded.NACE MR0103 (2005) limits the hardness of carbon steelcastings andplate to 200 HBW maximumand NACE MR0175(2003) limits the hardness to 22 HRC (237 HBW). Underboth specifications, carbon steel castings are acceptable in thenormalized and normalized and tempered conditions, andcarbon steel plate is acceptable in the hot-rolled condition aslongasthematerial meets themaximumhardness requirement.

The hardness of the carbon steel is largely controlled by thecarbon content, sokeepingthecarbon contentlow is essential tomeeting the maximum hardness requirement. Welding of thecarbon steel plates under NACE MR0103 (2005) is controlledby NACE MR0472 (2005) which also limits the maximumhardness in the base metal and the weld metal to 200 HBWwhile limiting the hardness of the HA Z to 248 HV 10. Themaximum hardness of 22 HRC is required in the base metal,weld metal, and HAZ under MR0175 (2003). The chemicalcomposition of the f iller metal used during welding must besimilar or identical to thebasemetal composition.Thehardnessmay beverifiedby theweldingprocedurequalification when allof the weldingparameters andfiller metal composition definedby the procedure qualification are controlled and followed. A

post weld heat treatment (PWHT) may be performed to ensurethat thehardnessvalues meet therequired specif ications. NACEMR0103 (2005) has a lower maximum hardness limit thanMR0175 (2003) to compensate for nonhomogeneity of someweld deposits and normal variations in production hardnesstesting using a portable Brinell tester.

During the last 15 years, compressor casings have becomelarger with increases in pressure rating. This has also beenapplied to applications where the compressor casings aresubjected to low temperatures. Design engineers now requirethicker sections dueto thehigher pressureratings to preventgasleakage at the various joints connections. For the thickersections, special controlled chemistry and heat treatments arerequired to achieve the desired toughness for low temperatureapplication. For low alloy nickel steels, theelements C, S, and P

are controlled to ensure good toughness is attained. Table 4showsresults of chemistries and their corresponding toughnessresults. Thesechemistriesnot only apply for steel platebut alsofor forgings andcastings. Sincetheapplicationof argonoxygendecarburization (AOD) and calcium-argon injection, cleanersteels, with sulfur as low as 0.005 percent, are attainable.Because of the addition of calcium compounds, the inclusionsthat remain resist elongation during rolli ng; remainingspherical. Consequently, inclusion shape control steel platesshow improved ultrasonic quality, macrocleanliness, andmechanical propertiesas can beobserved in Figures 11, 12, and13 (LukensFineline Steels).

Table 4. Results of Chemistri es and Their Corresponding Toughness Results.

Figure 11. Comparison of Sulfur Contents of 50 Heats of Conventionally

Processed Steel and 50 Heats of Lukens Fineli ne Double-O Five.

Figur e 12. Compari son of Charpy Impact Energy for C onventi onal and Fi neli ne ASTM A633C Plate, 4 Inches Thick, Normali zed.

Figur e 13. Compari son of Charpy I mpact Energy for ASTM A633C Plates, 1 to 2 I nches Thick, Nor malized, Transverse Ori entati on.




The introduction of inclusion shape control steel plates has alsominimizedthesusceptibility of lamellar tearingduringthefabricationof casings. Generally, this phenomenon can occur in thicknessesgreater than 3 to 4 inches. By achieving lower carbon, sulfur,phosphorus, and inclusion shape control structures, excellenttoughness properties can be achieved. However, for the largerthicknesses, acceleratedquenching during heattreatment using wateras the media from the austenitizing temperature is required. Tomaximize the quench affect of the water, the following is required:control of water temperature 80F maximum, good agitation of the

water, adequate amount of water for the weight of component (2gallons per pound of metal) andtimefromfurnaceto water less thanone minute. In some cases movement of the component in the watersideways or up and down is required (Figures 14, 15, 16, and 17)(Metals Handbook, 1981).

Figure 14. Surf ace Cooling Power of M oderately Agit ated Water Versus Water Temperature.

Figure 15. Effect of Concentration on Cooling Rate.

Figure 16. Effect of Temperature on Cooling Rate.

Figure 17. Effect of Agitation on Cooling Rate.

In Cameron’s (1989) paper, the changes in the 1987 ASME Boil er and Pressure Vessel Code Section 8 D ivi sion I asrequiredbyAPI 617 (2002) specified the required Charpy V-notch energyabsorption values had been madeafunction of theplatethickness.Refer toASME Boi ler & Pressure Vessel C ode - Secti on VII I (2007)to show the values required. Sincemost of thecompressor casingsarefabricated, the welding consumables haveto satisfy the impacttest requirement of the base material. Actual test results of theweldmentsand surroundingheat-affected zonedependstrongly onthe preheat, interpass temperature, thickness of individual weld

bead, heatinput rate, andpost weld heattreatment. Manufacturershave developed specific welding consumables, processes, andprocedures to obtain the required mechanical properties for thecasings. Since the soaking time of post weld heat treatments canaffect the toughness of the materials as well as the weldments,simulated PWHT is performed on the materials that are used forlow temperatureapplication.Figure18showseffectof PWHT timeon toughness(CharpyV-notch impact value).

Figur e 18. I mpact Energy Versus Temperature for SA516 Grade 60 after Vari ous PWHT.

Diaphragms

In the past, diaphragms were generally constructed from grey

cast-iron materials. When higher strength is required, ductileirons or fabricated mild steel is applied. In all cases, thecastingorfabrications areheat treatedtoproducelow levels of internal stressand difficulty with instability in service is virtually unknown. Inthe last 10years, the choice of diaphragmmaterial has been mildsteel with the blades either milled integral or welded and thetwo-piececonstructionboltedtogether.Oneof themain reasonsforthe change is that the mild steel materials can be repaired bywelding more readily as compared to grey cast-iron. The use of ductile iron that has improved weldability over grey cast-iron canalso beselected as amaterial choicefor diaphragms.

Sulphide Stress Cracking

There is extensive literature on sulfide cracking in oil well casingmaterials going back to about 1950. Incidents involving centrifugal

compressors, however, have been rare as reported by Kohut andMcGuire(1968) andMoller(1968). However,afailureduetosulphidestress cracking (SSC) can beserious; leading to loss of service of avital linkin theproduction chain of achemical orpetrochemical plant.ForSSC tooccur, thefollowingconditions must befulfilled:

•Hydrogen sulfide must bepresent.

•Water must bepresentin the liquid state.

• The pH must beacidic.

•A tensile stressmust bepresent.

•Material must bein a susceptible metallurgical condition.

195SELECTION OF MATERIALSAND MATERIAL

RELATED PROCESSES FOR CENTRIFUGAL COMPRESSORSAND STEAM TURBINESIN THE OIL AND PETROCHEMICAL INDUSTRY



When all the above conditions are present, sulf ide cracking canoccur with the passageof time.

As stated previously, all manufacturersfollow therequirementsspecified by NACE. In reviewing the work done Treseder andSwanson (1968) showed the effect of pH on sulphide stresscracking, Figure19 whereSc is an experimental stressvalue. ThisSc value increased substantially as the pH increased from2 to 5.Other work done by Keller and Cameron (1974) has shown theimportance of pH. In their work, quenched and tempered weldedAISI 4140 material stressed to 80 percent of yield strength failed

at pH 2.5, while at pH 4.2 no specimens failed. Theyield strengthof the base material was reported to be 126 ksi. From the samepaper the material AISI 4140quenched andtempered with a yieldstrength of 83 ksi was welded and given a temper treatment of 1100F. I n this case, at pH 2.5 all specimens failed. At pH 4.2, alarge percentage of specimens failed, while at 6.5 pH there wereno failures in those tests. It is important to note that these arecontrolled laboratory experiments. In practice, it is difficult tospecify a threshold concentration of H2S in a gas below whichsulfide cracking will not occur, because other components of thegascancausethepH tovary. Chlorides, for example, canlower thepH well below 4.3, whileother componentsmay causeanincrease.

Figure 19. Cri tical Stress for Sulfi de Cracking as a Function of pH.

Bolting

Boltingfor usein hydrogen sulf ideserviceenvironments is alsodefined by NACE MR0175 (2003) and MR0103 (2005). Bothspecifications allow the use of AI SI 4140 steel and AISI 410stainless steel at a maximum hardness of 22 HRC provided thematerial is quenched andtempered. BothAISI 4140andA ISI 410boltingis commonly available fromboltingsuppliers; however, anextratempering treatment is required to meet 22 HRC maximumhardnessrequirementandconformtoASTM A193GradeB7M forthe AI SI 4140 steel or ASTM A193 Grade B6M for AI SI 410stainless steel. For higher strength bolting materials, UNS 17400and UNS 15500 stainless steel may be used. NACE MR0175(2003) allows these alloys to be used at a hardness of 33 HRCwhile NACE MR0103 (2005) limits the hardness to 29 HRCmaximumfor pressure-retainingbolting.

Piping The piping of a compressor unit is typically manufactured from

carbonsteel piping. NACE MR0103(2005) limits the hardnessof carbon steel piping to 200 HBW while NACE MR0175 (2003)allowscarbonsteel pipingup to 22 HRC (237 HBW). Thecarboncontent of the carbon steel pipe has a large influence on thehardness, sothecarboncontentmust belimitedtocomply withthemaximum hardness limit. Under both NACE specifications, thepiping must also be thermally stress relieved following any colddeforming by rolling, cold forging, or any other manufacturingprocessthat results in apermanentouter fiber deformation greaterthan 5 percent. For carbon steel piping welds, NACE MR0103

(2005) defers to NACE RP0472 (2005) which allows a maximumhardness of 248 HV10 in the HAZ and a maximum hardness of 200 HBW in the base metal and weld metal. NACE MR0175(2003) requiresamaximumhardnessof 22HRC at all carbonsteelpipingweld locations.Whenmatchingfiller metals areused, apostweld heat treatment is not required if all hardnessvalues meet themaximum hardness requirement. The hardness values may beverified by the welding procedure qualification if all weldingparameters and the filler metal composition are defined by thequalification and are followed during production per NACE

MR0175 (2003). Under NACE MR0103 (2005), which defers toNACE RP0472 (2005), 5 percent of all production piping buttweldsmust behardness tested.

Spli tl ine Seals/O-Rings

Turbomachinery casings usually consist of two or morepieces,which areboltedtogether. With respect to centrifugal compressors,the seamcreated is sealed using elastomeric materials in the formof O-rings. Thechoiceof material used for this seal is vital to theoperation of the compressor, as it is typically the weakest link inthecasing in terms of chemical resistance, temperatureresistance,mechanical properties, longevity, etc. Due to the wide range of operating parameters encountered in a centrifugal compressorand wide range of reactions to those parameters by the variouselastomers, there is not a single, one stop choice. Table 5 is a

compatibility chart for some common gases andO-ringmaterials.In fact, choosing the proper material frequently consists of anevaluation and acceptance of several compromises, particularlyin cases with complex process gases. As a simplified example,fluoroelastomerswill work well in hydrogenandnitrogenseparatelyup to 300F. If they are mixed at 250F, such that a small amountof ammonia can form, the situation is no longer clear cut. Thefluoroelastomer is not compatible with ammonia, nitrile has amaximum temperature of 212F, and silicone is not compatiblewith hydrogen. Clearly acompromisemust bemadetoacceptwhatshortcoming will occurwhen thechosenmaterial is exposed totheoffending environmental parameter.

Table 5. Compatibili ty of Split Line Seal Materi als i n Various Environments.

Abradable and Rub Tolerant Seal Materials

Selection of abradable material for application in centrifugalcompressors is very important. There are a number of propertyfactors that need to be considered when selecting the abradablematerial:

•Abradability anderosion resistance•Compatibility of theabradable material with thegas

• Temperature limits of the abradablematerial

•Coefficient of thermal expansionof the material

Abradability and erosion resistance —In the paper presented byDowson, et al., (1991), abradablematerial mica-filledtrifluoroethanol(TFE), nickel graphite, andsiliconrubber werefoundtohavegoodtovery goodabradability. Theabradablesiliconaluminumpolyester,although not as good as those abradables stated previously, didperform well at lower rates of interaction and higher rubbing




velocities.Theauthor’scompany has applied thesiliconaluminumpolyester abradable material in various centrifugal compressorapplicationswith great success. Also since the 1991paper, nickelgraphite abradable seals have been used extensively in variousapplications for centrifugal compressors. However, to achieveoptimumabradability anderosion resistance, special control of thehardnessneedstobeachieved. Intheearly years of manufactureof theseseals, it was found that highheat input spray processes suchas plasma could reduce the percent graphite in the coating andthereby reduce the abradability. During testing of an abradable

nickel graphite seal the labyrinth rotating teeth were found to gallwith the seal causing excessive damage to the impeller seal eyelocation (Figure20).

Figure 20. Damage to Nickel Graphite Impeller Seal.

Compatibili ty of the abradable material with the gas —Themica-filled TFE is generally impervious to corrosive attackfrom most gaseous mixtures and would be acceptable with allhydrocarbongases, sour natural gas, andchlorinegaseousconditions.However, for wet gaseous conditions (water > 2 percent) certaindesign considerations must be addressed to account for waterabsorption leadingto possible swelling of theseal during service.

The aluminum silicone abradables would operate under the samegaseous conditions that are generally applied to conventional

aluminumseals. Exposed to extreme sour natural gases and chlorinegaseous conditions, these materials would be attacked severely. Thenickel graphitecoating could beused for all hydrocarbon gases andmost sour natural gases. However, for hydrocarbon gases thatcontain large quantities of carbon monoxide (CO), a reaction canoccur betweenthebasecoat of nickel andCO leadingtodelaminationof the nickel graphite abradable seal asshown in Figure 21.

Figure 21. Delamination of Nickel Graphite Seal Due to Reaction of CO with Ni.

Temperature limits of the abradable material —Thetemperaturelimits for thevarious abradablesareshownin Table 6.

Table 6. Temperature Limits.

Coefficient of thermal expansion of the materi al —Whendesigningseals, thecoefficientof thermal expansionof thematerialmust be taken into account at the design. The mica-filled TFEmaterial has a coefficient well above that of steel and, therefore,dimensional changes due to temperature must be calculated in theoverall design of the compressor. However, since the sprayedabradables are only 0.1 inch thick bonded to a metallic substrate,the coefficient of the substrate would apply for design purposes.For silicon rubber abradable material, thematerial is flexible/softand will deform elastically to accommodate any thermal strainscaused by thesubstrate.

Rub Tolerant Polymer Seals

Rub tolerant labyrinth polymer seals are seals with reducedclearances and, if contact is made betweenthestationary seal andsmooth rotating member, the stationary teeth will deflect duringcontact withoutwear or damageto the rotor or seals.

The rub tolerant plastic seals that are used in centrifugalcompressors today are the new thermoplastics, which havebetter resistance to elevated temperatures. The thermoplasticsmatrix materials aretougher andoffer thepotential of improvedhot/wet resistance.

Becauseof their highstrainstofailure, they aretheonly matrixesthatoffer the new intermediate modulus, high strength (and strain)carbonfiber tousetheir full strain potential in thecomposite. Thesematerials include such resins as polyetheretherketone (PEEK),which is intended to maintain thermoplastic character in thefinal composite. Others, such as polyamideimide (PAI), which isoriginally molded as a thermoplastic and is then postcured in thefinal composite to produce partial thermosetting characteristics.

The partial thermosetting characteristic of the PAI enables animproved subsequenttemperatureresistance(Engineered Materials

Handbook: Composites, 1987).

When considering thermoplastic materials in rotating equipment,one has to understand their thermal properties (Table 7). The twothermoplastic materials used exclusively as labyrinth seals incentrifugal compressors are PEEK with additives and PAI. For thePEEK materials as labyrinthseals the temperaturelimit is dependanton the glass transition temperatures Tg of the material. Addition of additives such as chopped or continuous wound carbon f ibers orgraphitepowder orPTFE will not increasetheTg of thematerial. The

Tg is the temperature at which the crystalline polymer changes to aviscous or rubbery condition. In other words, the material has adramatic change in properties. Generally, for labyrinth seals fromPEEK material, theoperation temperatureis limitedto 290F.



197



Table 7. Properti es of Thermoplasti cs.

For PAI materials, since the material is partial crystalline incharacter with some amorphous features, the Tg of the material ishigher. However, when the crystalline character of the polymer isdecreased, its resistance to solvents and water decreases also. Thehigher the degree of crystallinity, the higher the modulus and thehigher theresistanceto solventsandwater. In thecaseof PAI carefulconsideration must be given to attack from amines, ammonia,oxidizing acids, and stray bases. Due to its partial crystallinity incharacter, PAI is prone to moisture absorption. Due to the moistureabsorption, dimensional changes can occur that need to beaddressedat the design stage and in manufacturing/storage. Generally,thermoplastic resin suppliers provide some information on waterabsorption after a 24hour immersion in water. A more severetest isthe24hour boilingwater test.

Duetothehighthermal expansioncoefficient of both thePEEK andPAI materials, accuratecalculationsof the growthof the sealswith respect tothediaphragmsandrotatingcomponents needtobedoneto enable one to calculate the clearances.

Also,in themanufactureprocessingof PAI or PEEK intoatubularform, thereexists a sizelimitationfor polymer labyrinth seals.

•PAI—In tubular bulk form the size limitation is 35 inches.However, larger sizes can be constructed up to 45 inches bysegmentingtheseals.

•PEEK with chopped carbon fibers—Generally supplied inmolded bulk formwitha sizelimitationof 30 inches.

•PEEK with 68 percent continuous wound carbon fiber—Generally suppliedin atubularformandthereis nosizelimitation.

Abradable seals have been used successfully in centrifugal

compressorsfor reducingclearancesandimprovingtheefficiency of compressors. Theauthor’s company has appliedabradable materialsuch as Ni graphite, aluminum silicon polyesters, and fluorosint tonumerouslabyrinthseal applications with greatsuccess.

Careful considerationmust beappliedtoensurethatrubtolerantpolymer seals can beutilizedin centrifugal compressor labyrinths.

Thetests that weredoneat theauthor’s company indicated that thePAI material for labyrinthseals may not besuitable for temperaturesgreater than 100F using similar clearances to that of abradables.

The PAI material at a temperature of 150F was found to weardramatically when it came into contact with the rotating member.Other polymer materials, such as PEEK or carbon wound PEEK,may bemoresuitable (Dowson, et al., 2004).

STEAM TURBINES

Steam turbines present some different problems relative tocentrifugal compressors. Temperatures are usually higher, whichaffectsthevarious temperaturedependentprocesses likecorrosionrate and degradation mechanisms, such as creep, must now beconsidered. The key components of a steam turbine are the rotor,blades (buckets), casing, andbolting.

With the exception of 12%Cr blading alloys, low alloy materialshave been conventionally used for the major components of steamturbines for many years. For example, 1CrMoV has been used forhightemperaturerotor forgingsand3½NiCrMoV for lowtemperatureforgings. High temperature pipe work and castings for valve chestsandturbinecasings utilize 2¼CrMo or CrMoV. Thesematerials were

available in the 1950s and 1960s and over the past 30 to 40 yearsmaterials development concentrated on optimizing the properties of these materials for their application to larger components, which inturn, increases thereliability of components andtheir fabrications.

Rotors

Mechanical drive steam turbine rotors are manufactured fromlow alloy steel forgings. Rotors consist of a shaft and a series of disks, which may be monoblock, in which the disks are integralpartsof theshaft (A type), disksshrunk fit ontotheshaft (B type),

or sections of the rotor may be welded together (C type).Mechanical drive turbines are in general “A” or “B” type rotors,withthe majority being“A” type.

For the higher temperature application up to 1025F, the1Cr1Mo¼V steels are currently in wide use but are limited to1025F. For higher temperatures, the 12%Cr steel forgings havegoodexperienceat1050F,wherethey haveanexcellentcombinationof creep strength, ductility, and toughness. However, there is adrive to use the most creep resistant version of 12%Cr up totemperaturesof 1100F. Inorder toprovideanadditional margin of safety, theallowablecreepstrengthat1100F needstobeequivalenttothat currently usedwith12%Cr steel at 1050F (Gold andJaffee,1984; Armor, et al., 1984). For developing improved alloys, atentativegoal for rupture strength has been established at 14.5 ksifor a rupture life of 100,000 hours at 1100F. Figure 22 shows a

comparisonof 100,000hourrupturestrengthsfor variouslow alloysteels, 12%Cr and other development rotor steels. The Figure 23shows the Larson-Miller rupture curves for commercial anddevelopmental 12 percent rotor steel (Newhouse, 1987). Othercritical material properties for rotor integrity are toughness,resistance to crack, initiation under creep, thermal fatigue andresistance to subcritical crack propagation in creep and fatigue.

Toughness is discussed in more detail below.

Figur e 22. Plot Showing 100,000 Hour Rupture Strength of Vari ous Rotor Steels.

Figure 23. L arson-Mil ler Rupture Curves for Commercial and Developmental 12 Percent Rotor Steel.




Toughness

In CrMoV rotors, ASTM A470 Class 8 specification limits thefracture appearance transition temperature (FATT) to 250F. TheFATT values for the base material normally range from 185 to260F. Since theadvancementof fracturemechanics technology, ithas now become possible to characterize toughness in terms of acritical crack size ac. The cold start sequence of a rotor is mostcritical andthecritical flaw sizeac basedonanalysis of temperatureandstresscanleadtocatastrophic failure(ViswanathanandJaffee,1983). For the Gallatin rotor, the temperature at the time of the

peak stress74ksi was 270F. Thecombined crack sizereached itslowestvalue0.27 inches at this temperatureandstress(Figure24).

Figure 24. Il lustration of Cold-Start Sequence and Associated Variations in Stress ( σ ), Temperature (T), and Cr it ical Fl aw Size (a c ) as F uncti ons of Time f rom Star t.

Variations in temperature, stress and material in homogeneitythroughout therotor dictatethecritical flaw sizeac. By usinglowerscatterband values of K IC, the critical flaw size can be converted(Schwant and Timo, 1985). Another method of estimating thelower-bound values of K IC for CrMoV steel is by using theexpression (J ones, 1972):

where:K IC is expressed in MPa

T is in C

Although equipment manufacturers have records of the FATTvalue of the rotor material prior to service, the effect of temperembrittlement duringserviceincreasestheFATT valueanddecreasestheK IC value (Figure25) (Viswanathan andJaffee, 1983).

Figure 25. Effect of Temper Embrittlement on Fracture Toughness of CrMoV Rotor Steel.

Themaximumtemper embrittlement occurringat locationof therotor exposed to temperatures from 700F to 800F. Consequently,evaluation of FATT (or K IC at the concerned location) in theservice exposed condition is critical for damage assessment(Dowson, et al., 2005).

An ASTM special taskforceonlargeturbinegenerator rotorsof subcommitteeVI of ASTM CommitteeA-1 on steel has conducteda systematic study of the isothermal embrittlement at 750F of vacuumcarbondeoxidized(VCD) NiCrMoV rotor steels. Elements,suchasP, Sn,As,Sb, andMo werevariedinacontrolledfashionandthe shifts in FATT, () FATT were measured after 10,000 hours of exposure. Fromtheresults, thefollowingcorrelationswereobservedin Equation(2):

where:FATT is expressedin degreesFahrenheit andthecorrelationof allthe elements are expressed in weight percent. According to thiscorrelation, theelementsP, Sn,andAsincreasetemper embrittlementof steels, while Mo, P, and Sn interaction decrease the temperembrittlement susceptibility.

All available 10,000 hour embrittlement data are plotted inFigure 26 as a function of calculated FATT using Equation (2)(Newhouse, et al., 1972). A good correlation is observed betweencalculated and experimental FATT. The scatter for these data isapproximately ± 30F for 750F exposure and ±15F for the650F exposure.

Figure 26. Cor relation Between Compositi onal Parameter “ N” and the Shift in FATT of NiCrMoV Steels Following Exposure at 650 F and 750 F for 8800 Hours.

Other correlations for determining the temper embrittlementsusceptibility of steel, such as the J factor proposed by Watanabeand Murakami (1981) and factor proposed by Bruscato (1970),arewidely used.Thesefactorsaregiven by:

The Figures 27 and 28 show relationship between increaseof FATT and J factor and factor at 750.2F (399C) for a3.5%NiCrMoV steel.



199

m



Figure 27. Correlation Between Compositi onal Factor “ J” and the Shift i n FATT of NiC rM oV Steels Followi ng Exposure at 650 F and 750 F for 8800 Hours.

Figure 28. Relati onship Between I ncrease of FATT and .

In the high temperature regions, creep and creep rupture are aconcern. Thetraditional approachis to usea Larson-Miller plot asshown in Figure 29. The design stresses are generally based onthe 105 hours smooth bar creep rupture stress divided by someappropriatesafety factors.

Figure 29. Larson-Miller Stress Rupture Curve for 1Cr 1Mo ¼V Rotor Steel.

Steam Turbine Blades (Buckets)

Therearebasically three groups of steamturbine blade materialusedby turbinemanufacturers.Thesearevarious grades of 12to13percent chromium (cr) steels with additions of Mo, W, Cb, andV,higher chromiumprecipitation hardeningsteels suchas17-4PH andtitanium alloys. Table 8 gives a listing of the commercial availablematerials used in blading (ASTM A1028, 2003). Additional dataregarding the heat treatment and mechanical properties can befound in Aerospace Str uctural Metals Handbook: Volume 2 (1988),andBriggs andParker (1965) (Figures 30 and31).

Table 8. Composition and Mechanical Properties of Commercial Blading Materials.

Figure 30. Rotational Bending Fatigue Behavior of Type 403 at Vari ous Temperatures.

Figur e 31. Stress-Range Di agrams at Vari ous Temperatures for Type 403 Heat Treated to 26-32 HRC.




In designing blades one must take into account the materialproperties, the operating conditions, and the quality/purity of thesteam. The principal failure mechanisms that occur in blades arehigh cyclic failure and creep rupture. Blades are designed toprevent creep rupture failures; therefore, it is rare that steamturbine blades fail by this mechanism. In general, most bladefailures are due to high cycle fatigue caused by a number of factors. Some of the factors arelisted below:

•Dynamic stresses caused by nonsteady steam forces, nozzlewakes, thermal transientper revolution diaphragmharmonics, and

flow instabilities

•Strongexcitingharmonics of rotational speedsuchas thenozzlepassing frequency

•Steam purity and its effect on corrosion in fatigue and pittingcorrosion

Speidel (1981) highlighted the satisfactory corrosion fatigueresistance of X21CrMoV121 steel in a good purity steam(Figure 32). However, the growth of fatigue cracks in 12percentchromiumsteels is greatly enhanced by thepresenceof chloridesolutions at low cyclic stressintensities andhigh meanloads(Figure 33). Speidel (1981) also illustrated how theK th(threshold stressintensity) is greatly reduced by thepresenceof certain environments (Figure 34). Other work done by Batte

and Murphy (1981) showed similar results on the fatiguestrength reduction in various aqueous solutions (Figure 35).With thesedata, it can beconcluded that the most concerningconditions for fatigue crack growth is a high mean stress withsuperimposed cyclic stresses in the presence of aqueoussolutions. These conditions are the service conditions thatsteam turbine materials are subjected to. It is well known thatshot peening enhances the fatigue resistance of metallicmaterials by introducing compressive stresses at the surfacelayer of the component. However, in a corrosive environment,pits may penetrate the compressive stress layer, therebynegating the benefit gained by peening (Figure 36). Theimportance of steam purity and corrosion resistance of theblade material are clear when considering that it is not feasiblein most cases to avoid the stressstateandexposureto aqueous

solutions present in the turbine and that treatments such aspeening are not reliable.

Figure 32. Fatigue and Corrosion Fatigue of X21CrMoV 121 in Air, Deaerated Water, and Aerated Hot Chlor ide Soluti ons.

Figure 33. Effect of Chloride Solutions on the Fatigue Crack Growth Rate of 12%Cr Steels.

Figur e 34. Effect of Envir onment on the Threshold Stress Intensit y K th of 12%Cr Steels.

Figur e 35. Effect of Vari ousAqueousEnvironmentson Fatigue Strength.





Figure 36. Effect of Envir onment on Fati gue Resistance of Shot Peened 12%Cr Steel.

Casings

The various composition of steels used for turbine casings are

shownin Table 9. Duringtheyears thedrivetowardimprovingcreepstrengthto accommodatethesteadily increasingtemperatures ledtoprogressive changes in material fromthe C-½ Mo and1Cr½Mo tothe 1Cr1Mo¼V and 2¼Cr1Mo. In the late 1960s and early 1970s,numerous instances of reheat cracking (stress relief cracking) inthe weld heat affected zones occurred in the 1Cr1Mo¼V and theimportance of rupture ductility was realized. The material½Cr½Mo¼V was standardized by some manufacturers whoimplemented stringent specifications relating to control of residualelements(particularly phosphorus, antimony, tin, copper, aluminum,and sulfur), deoxidation practices, and welding procedures. Casingdesigns were modified to eliminate manufacturing and in-servicereheat cracking. Other manufacturers utilized the 2¼Cr1Mo steeldueto its higher creep ductility, higher low cycle fatigueresistance,andbetter weldability. The current designs use either ½Cr½Mo¼V

or the 2¼Cr1Mo steel material. For steam temperatures of 1050

Fand above, and up to 1100F, the 9 to 12%Cr casting grade steelswill be utilized.

Table 9. Vari ous Composit ion of Steels Used for Turbine Casings.

The optimization of low alloy materials has taken place over aperiod of 50yearsormore.Thedevelopmentandoptimizationof the9 to12%Cr haveonly just begun. A lot of work remainstobe doneto characterize these materials for design lives of up to 250,000hours and to cast variation in their properties. The development of

joints between the modified 9%CrMo and low alloy material hasbeen done in order to optimize selection of welding consumables,welding parameters, and post weld heat treatment. Creep tests onsuch joints have indicated that in long-term tests, their rupturestrengthfalls near thelower boundof thelow alloy parentmaterial’sstrength. Due to composition gradients, carbon depleted, ferriticzones can formon the low alloy side of the interface between thehigh and low alloy materials. Consequently, joint locations aredesigned suchthat thetemperatureandstressarelowersothat creepand carbondiffusion in servicewill be very limited.

Nozzles and Di aphragms

Insteamturbines, stationarynozzlesanddiaphragmsareselectedbaseduponstress/temperature, oxidation,andcorrosion. Thebladesof nozzles and diaphragms are made from wrought 13%Cr seriesstainless material. The blade holders are manufactured from mildsteel, ductile iron, or low alloy steels. The choice of materialdepends on the method of construction and the operating designstressat temperature.

Duringservice,erosion duetowet steamcan occur onthelatterstages of thesteamturbinediaphragms. Normally repairs in those

areas are done by depositing weld material that has improvedwater/steam erosion resistance, i.e., AWS E309/ER309 orInconel® weld consumable. In some instances, a mechanical fixmay be performed using an austenitic stainless or nickel basewrought material.

High Temperature Bolting

In a steam turbine, bolts in flange joints operating in the creeprange at temperatures up to 1050F must be able to withstandsteampressures upto2 ksi or in somecases even higher.Themainrequirementof bolts in thecreeprangeis tomaintain jointswithoutrelaxing below the required design stress limit, which may allowleakage. Consequently, theimportantproperty for boltsis thestressrelaxationcharacteristicsof thematerial. Foragivenjoint, theloadrequired to exceed the steam load is applied to the flange area by

tensile loading of the bolts. During service, creep in bolt causesrelaxationof this initial load. Theelastic strainsproducedby initialtighteningof thebolts areprogressively converted to creep strains;elongating the bolt and thereby, reducing the effective load in the

joint. For design, the final relaxation stresses must be in excess of thedesignstressto keep thejointtight. Depending onthematerialandduty requirements, bolts are usually tightened to a predefinedcold strain, e.g., 0.15 percent fromwhich the initial stress on thebolt can becalculated. Consequently, the bolts need to retain theirdesign stress between each overhaul when the bolts are removedfor maintenance of the machineandretightened on reassembly.

Typical composition of bolt steel, their properties, and variousnational standards are summarized inTables10and11(Everson,et al., 1988). For convenience, the Central Electricity QuenchingBoard of the United Kingdom has classif ied thesecompositions

into grouping numbered one through eight. Several of thesegroups of materials show stress relaxation behavior after 30,000hours as a function of temperature when subjected to a coldprestrain of 0.15 percent (Figure 37) (Branch, et al., 1973). Inselection of bolt material, the compatibility of the thermalexpansion coefficient of the bolts with respect to the jointmaterials, as well as its susceptibility to various fracturemechanisms, must betaken into account. When sufficient stressrelaxation data are not available, one can apply the usefulcorrelations that exist between the creep rupture data andrelaxation stressat 0.15 percentstrain (Figure38).

Table 10. Various National Bolting Material Specifications.

Table 11. United Kingdom Bolting Material Specifications.




Figure 37. Comparison of Stress-Relaxation Behavior after 30,000 Hours as a Function of Temperature for Various Bolt Materials Subjected to a Cold Prestrain of 0.15 Percent.

Figure 38. Relationship Between Relaxed Stress and Rupture Strength at the Same D uration for Times from 1000 to 30,000 Hours and Temperature from 885 to 1110 F (475 to 600 C).

Fracture of a bolt can occur if the local creep strain reaches thecreep ductility of the material from which the bolt was made.Consequently, low rupture ductilities lead to notch sensitive boltmaterial. Numerous failuresof 1Cr-1Mo-V steel bolts wereattributedto notch sensitive failures. By reducingthe impurity elements of thebolt steel, appreciable improvements have been made to ruptureductility without compromising the creep strength (Figure 39). Theresidual element content can be reduced by careful scrap selection,avoidanceof air melt, anduseof double vacuummelting(Figure40).

Figure 39. Relati onship Between Residual Element Content and Ductility.

Figure 40. Effect of Double Vacuum Melting on Stress-Rupture Properties of 1Cr-Mo-V Steels at 1020 F.

Other improvements in notch ductility were in a new class of CrMoV containing titanium and boron with subsequent grainrefinement. Further improvements in the rupture ductility of 1CrMoVTiB steel havebeenmadeby reducingthemajor embrittling

elements. The application of very clean steel practices have shownbeneficial effectsonductility andnotchstrength. This material willcontinue to beused as a very cost-effective option for temperaturesup to 1049F (565C). Also nickel-based super alloys such asNimonic 80A and IN901 have been used for some of the hightemperature bolts wherestressrelaxation is an issue.

REPAIRS

Intherotating equipmentindustry, serviceof componentsrequireseither replacement with new or refurbishment of the components.

The refurbishment can mean restore the component to its originaldimension either bywelding/mechanical fix or other sprayed/platingprocesses. All OEMs have developed repair procedures for most of the long lead delivery components, e.g., components such asimpeller, shaft, casing, anddiaphragms.API RP 687(2001) specifies

the minimumrequirements for performing therepairs to rotors. Throughout the last 20 years hundreds of rotors and shafts for

both compressors andsteam turbines have beensuccessfully weldrepairedby variousOEMs.Weldedrotor restoration, likeany othercritical repair technology, requires a highly analytical approach toassure component and machine reliability. Most OEMs developproperty datafor weldmentsandheataffectedzonesfor thevariousmaterials.This would includematerialsfor bothsteamturbinesandcentrifugal compressors. The data that are generated, but are notlimited, areas follows:

•Roomandelevated temperaturetensileproperties

• Impact and fractureappearancetransition temperature data

•Creep/stress rupture data

•Fatigueproperties•Stresscorrosioncracking(SCC) threshold limits

Several papers (LaFave, 1991; Dowson, 1995; Dowson andWiegand, 1996) outlinesomeof thematerial property datathat canbegenerated. Figures 41, 42, and43showsomegeneric examplesof datathat aretypically used. Thesedataallow for theweld metalproperties to becompared against mechanical analysis results andcompany designcriteria to assure that the proposed repair will bereliable. Shorter termtesting such as weldability, tensile hardness,andtoughnessis routinely doneon thebasemetal, weld metal, andHAZ for eachengineered rotor repair.





Figure 41. Typical Larson-Mi ll er Curve Generated from Vari ous Weldments.

Figure 42. Typical S-N Curve from Fatigue Testing of Weldments,Fully Reversed (R = 1), Endurance Limit Set to 10 8 .

Figure 43. Stress Corrosion Crack Growth Rate as a Function of Stress Intensity of Weld Metal, HAZ, and Base Metal.

When repairs to impellers are required, it is generally due tosome form of mechanical damageor corrosion/erosion attack. If the impeller has failed by fatigue, the component is generallyreplaced. When corrosion attack has occurred, repairs will bemade by applying corrosive resistant welding consumables. Forerosionattack, various coatings can beapplied to extend the lifeof the impeller.

COATINGSDespite best effortsin choosing a suitable alloy, it is frequently

impossible to find a single material that is ideal for the particularapplication. In this case, compositesystems, i.e., coatings, may beused to exploit the properties of two or more materials. There areseveral reasons why coatingsmay beused, rangingfromenhancinga particular property for a specific applicationto reducing costbyallowingfor theuseof lessexpensivesubstratematerials. Coatingsare also often designed to bemultifunctional, addressing multipleproblemareassimultaneously.Thepetrochemical industry presentsseveral opportunitiesfor theusageof coatingsduetothefrequentlyharshconditions encountered in this service.

Compressors

Corrosion

Corrosion is acommonproblemin several industries, includingturbomachinery.It isanelectrochemical attack that canoccur asaneven attack of the surface (general corrosion) or uneven attack(pitting), as well as lead to stresscorrosioncracking. Coatings forcorrosion protectionwork by acting as a barrier that separates thesubstrate material and the environment. Typically, the coatingmaterial is viewed as a sacrificial layer and has slower reaction

kinetics thanthebasemetal in theparticular environment.For selectionof materials for sour gas service, API 617 (2002)

refers to NACE MR0103 (2005) and MR0175 (2003), whichallow theuseof coatings for general corrosionresistance, but notfor protection againststresscorrosion-cracking.This is duetothefact that localized defects in the coating can lead to a stresscorrosion crack. Withany coating, thereis always a risk that thecoating will be removed at a small location, either by erosion,localized pitting, orforeignobject damage, which will exposethesubstratetothecorrosiveenvironment. Diffusion coatings aretoothin to providea completebarrier between the substrateand theenvironment. Metallic, electrolytic coatings contain microscopiccracks that will allow the H2S to contact the substrate material.Hydrogen sulfide can diffuse through polymer coatings to reachthesubstratewhilethermal spray coatings areporous andwill not

provide adequateprotection. There appearsto be a limited number of coatings advertised bythe various OEMs that areapplied solely for corrosionprotection.Part of this is duetothefact that changingthebasemetal toamorecorrosion resistant material, such as a stainless steel, often givessatisfactory results. A second reason for the lack of corrosioncoatings is that corrosion is often accompanied by fouling incentrifugal compressors and therefore the antifoulant coatings,discussed below in more detail, typically are designed to preventcorrosionas well. Onecoatingthat theauthors’ company has usedwithsomesuccessis abakedphenolic coating. This coatingis usedin certain applicationswhereit performs better than stainless steelor where the increased price of using stainless steel cannot be

justified by the end user.

Fouling

Fouling is a common problem in compressors and to someextent, steam turbines. Fouling refers to the build of solids,usually polymeric materials, ontheinternal aerodynamic surfacesof the machine. While it does not usually lead to catastrophicfailure, it does gradually reducethe efficiency of the machinebyincreasing the mass of the rotor, altering the aerodynamics, andblocking flow paths. If left unchecked, foulingcan block theflowpath to the extentthat productionis stopped or cause imbalancesthat can damagethe machine. Depending on the service, foulingsubstances may come from outside of the machine or begenerated internally. External foulants may come from airbornesalt, submicron dirt, and organic or inorganic pollutants in theprocess gas. A well-maintained filtration systemusually helps tominimize this typeof fouling(Meher-Homji, et al., 1989; Guineeand Lamza, 1995). In petrochemical compressors, the situation

is much more complicated, as the foulants can be generatedinternally. For example, in ethylene cracked gas compression,fouling results fromthe polymerization reactions intrinsic to thecompression process. Fouling has imposed significant cost onpetrochemical production.

A material that is required to resist fouling must have excellentreleaseproperties. Materials with acombinationof low coefficientof friction andchemical inertnessareusually used for this application.A common and widely known coating material for centrifugalcompressors is polytetrafluoroethylene (PTFE [Teflon®]). These aremulticomponent, sprayed coatings designed for fouling andcorrosionresistance. Wang,et al. (2003), showedadramatic decrease




in timerequiredtoreleaseanapplied foulant onsamplescoatedwithtwo PTFE typecoatingsoffered bytheauthor’scompany versus baresteel samples (compare E and E+versus steel in Figure 44). Figure45 shows an example of a compressor rotor with PTFE coatedimpellers. Unfortunately, PTFE coatings are removed by erosiveliquids (example water washing) or solids. Electroless nickel (EN)has also been shown by Dowson (2007) to exhibit excellent releaseproperties (523 in Figure 44), while remaining adherent in erosiveconditions. Infact, thereleasepropertiesareasgoodasor better thanresults from PTFE. EN is applied by submerging the part into a Ni

and P containing solution where an autocatalytic process plates thepart witha well-bonded, amorphous Ni-P alloy. TheP in thealloy isbelieved to be responsible for the release properties, whileamorphousnatureaidsin corrosion resistance.

Figure 44. Comparison of Fouling-Release Performance of Bare Steel and Coated Panels.

Figure 45. Centrifugal Compressor Rotor Coated wi th a PTFE Type Coating.

Case Study of Electroless Nickel Coating

The application of electroless nickel as an antifoulant coatingis a relatively recent advance. A brief survey of antifoulantcoating offerings by OEMs showsthat themajority do notoffer anelectrolessnickel coating. Due to the current lack of offerings, anupdateof a case study initially introduced by Wang, et al. (2003),is presented below in order to justify the useof electroless nickelas a corrosionandfoulantresistantcoatingoption.

Background A major chemical plant in Corunnahas a centrifugal compressor

that usedtobecoatedwiththecoating3P for antifouling. Thecoatinghad suffered severe deterioration two times during the four yearoperation since 1997. Analysis indicated that the deterioration wasrelated to the heavy washing injection, which contained aggressivechemical additives, and steam-cleaning operation. However, thewashing and cleaning were essential to the plant because of theextensive fouling and efficiency drop. In order to withstand theinjections, the compressor rotor and diaphragms were recoated withEN during a plant turnaround in September 2001. The compressorhad been in service until 2006with satisfactory performance.

Operating Status of the Electroless Nickel Coating

The compressor with EN coating had been successfully operatedsince the 2001 turnaround until 2006. The rotor vibration has beenmonitored at a much lower level than the previous run periods,as shown in Figure 46, which indicates that the fouling has beeneffectively controlled by the EN coating. This improvement isattributed to the ability of EN to withstand the heavy washingoperation as compared to the 3P coating. The oil washing andchemical injections have been kept in the same manner as in theprevious operatingperiods. TheCorunna plant feels that the EN has

successfully functioned as an antifoulingcoating. When the unit wasreplaced in 2006 with a new machine, the customer requested thesamesuccessful EN coatingto beapplied to therotor andinternals.

Figure 46. Vibration Data Before and After Application of Electroless Nickel to t he Rotor.

After theoriginal unit waspulledout of service, examinationof therotor shows that the EN coating remained in tact after the fiveyearoperational period. Pictures of acoated impeller aregiven in Figures47 and 48. The coated impeller is free of corrosion and foulantbuildup, and the impeller serial number remains readable. This isproof of theremarkableantifoulant propertiesof theEN coating.

Figure 47. Picture of Fir st Stage I mpell er with Electroless Nickel Coating after 5 Years of Service. The Coating Remains Intact and the Seri al Number Is Cl early Visible.



205



Figure 48. Pi cture of Fir st Stage Impell er with Electroless Nickel Coating after 5 Years of Service. The Gas Path of the Impeller Has No Foulant Buildup.

Erosion

Erosion can be a significant issue in some applications of centrifugal compressors. It can occur as solid particle or liquid

dropleterosion. Solid particle erosion is usually causedby externalcontaminants, such as those mentioned above. Liquid dropleterosion can be due to the condensation during compression orintentional injection of liquid for cleaning. The particle or dropletis accelerated by the carrier gas. The resultant impact upon themetallic compressor components removes small amounts of metal; creating pits and microcracks at the surface. Typically, thecompressor designis robust andthe material removal rateis suchthat erosion on its own is not a major problem. Problems arisewhenerosion is combinedwith other factorssuchas corrosionandcyclicstresses. Incorrosiveenvironments, erosioncancompromiseprotection schemes andcause an accelerated corrosion attack. Anexample of this was mentioned above, where erosion can removesome antifoulant coatings. When cyclic stresses are present, themicrocracks created can easily serveas initiation sites for fatigue

cracking. Figure49showstheleadingedgeandfracturesurfaceof a compressor impeller blade that failed by fatigue initiated bydamagefromliquid droplet erosion.

Figure 49. Stereomicrograph Showing a Centri fugal Compressor Impell er Blade that Failed Due to Li quid Dr oplet Erosion.

Erosion can often be thought of as a type of wear; therefore,similar coatingsareusedtopreventboth. Coatingsusedfor erosionprotection areusually hardcoatings, suchas tungstenor chromiumcarbide. Any surface exposed to the gas path should be coated inapplications where erosion is a problem. One problem that is anissue withall coatings for compressors, but is particularly true for

thesehard, erosion coatings, is the methodof application. Carbidecoatings are typically applied using techniques such as highvelocity oxyfuel (HVOF), plasma spraying, or detonation gun.

These processes have certain requirements, such as spray distanceor line-of-sight, that may not be possible for the gas paths of smaller, closed impellers or other parts with poor accessibility.

Steam Turbines

Corrosion

Under ideal conditions, corrosion is not a major issue in steamturbines. Stainless steel materials perform extremely well withoutcoatings in the elevated temperature, steam environment. Thecorrosion product formed on the stainless steel is usually a thin,uniformlayer that grows slowly with time. Corrosion becomes aproblem when there is an underlying problem with the process.Steam purity is the most common problem. Impure steam carrieswith it elements that increasethe corrosion process. Typically, theamountof impuritiesin thesteamis not largeenoughtochangethegeneral corrosionrate; rather they causelocalizedattack thatleadstopitting.Thesepits provideeasyinitiationsitesfor fatiguecracks,asshownin Figure50. Figure51isaplotof stressamplitudeversuscycles to failure by fatigue. The drastic reduction in stressamplitude required for a given fatigue lifetime illustrates the easein which fatigue cracksformin corrosion pits.

Figure 50. Micrograph Showing a Crack Initiating from a Corrosion Pit.

Figure 51. Plot Showing the Effect of Corrosion and Pitting on Fatigue Life.

Similar to compressors, there are not many coatings applied tosteamturbinecomponentsfor corrosion only. If thereareproblemswith corrosion, usually other problems are occurring, such aserosion, which also must beaddressed by a coating. The authors’company has used chromium diffusion coatings in the past wherepitting corrosion was a problem.

Erosion

Erosion is a serious problemin steam turbines. Thegas path of asteamturbineismuchmoreclosedthanthat foundin acentrifugalcompressor, providing more area for erosive media to impact.




Erosion-corrosion mechanisms are more prevalent due to thehigher temperatures. The chances of cyclic stresses are also veryhighdue to the complicated stressstatefound in turbine blades.

Solid Particle Erosion

Theerosionoccurs ontheblade-vaneleadingedgescausedby theexfoliation of scales from the boiler tubes mainly during transientconditions. For rotating blades, sprayed Cr3C2 coatings are appliedusing processes such as detonation gun, plasma, and high velocityoxygenfuel processesto protect againstthescales. Other processes

such as diffused boride coatings have successfully been applied tostationary nozzles. Figure 52 shows a photomicrograph of boridecoatingonAISI 422. Theboridediffusion coatings appliedby packcementation andthe Cr3C2 coatings applied by the spray processesmentioned abovewill continue to bethe industry bestchoices.

Figure 52. Cross-Section of a Bori de Di ffusion Coating on AISI 422.

Liquid Droplet Erosion

The damage of rotor blades in the latter low pressure stages of steam turbines by condensed water droplets can bea problem thathas troubled designers and operators for many years. The damageconsistsof removal of material from the leadingedge andadjacentconvex surfaces of the moving blade. It is related to the wetnessconditions in the low pressureregions and the velocity with whichthesurfaceof thebladestrikesthewater droplets.Stressesproducedby impact of drops havebeen calculated by existing theories to be

sufficiently high to initiate damage in the latter stages of blades ina steam turbine. Turbine experience and laboratory testing haveshown that erosion rate is time dependent with three successivezones: a primary zone in which damage is initiated at slip planeswithlittle or no weight loss, a secondary zone where the raterisestoamaximum, andf inally atertiary zonewheretheratediminishesto a steady-state value. Moreover, it is this tertiary region that isimportant to designer and operator alike rather than the initial andsecondary zones since it is this region that the turbine erosionshieldsoperatefor most of their lives (Figures 53and54).

Figure 53. Erosion Rate of 630 D PH, 18W-6Cr-0.7C Tool Steel Comparator Specimens.

Figure 54. I mages Showing Erosion of Blades from the Same Turbine after 400 Hours and 70,000 Hours.

While not strictly coatings, various countermeasures have beenutilized by OEMs, including water drainage devices in a cylinderwall, flameorinductionhardeningof theleadingedgeof theblade,and applying stellite or tool steel strips to the leading edge of ablade by welding or brazing. All of theabovemethods showsomedegree of success, with the stellite material providing the better

performance in the morestringent water droplet environment.Fouling

Fouling and corrosion can also be a problemin steam turbinesnotonly causingmaterial damagebutalsocangradually reducetheefficiency of theturbine. Industrial turbines, whether condensingor noncondensing, can encounter problems with deposits buildingup on the turbine airfoils. In a turbine, hydroscopic salts, such assodium hydroxide, can absorb moisture when superheated steambecomes saturated and condenses in the latter stages of theturbine/Wilson line. Wet sodium hydroxide has a tendency toadhere to turbine metal surfaces and can entrap other impuritiessuch as silica, metal oxides, and phosphates. Once these depositshave formed they can be difficult to remove. Build up of thesedeposits may be a causeof decreasein efficiency and possibly an

increasein vibration. A smooth clean steam path will not collectdeposits so easily as a dirty, previously contaminated surface.Consequently, a previously contaminated turbine will accumulatedeposits more rapidly than a clean one. Therefore, it is desirableto prevent further deposit buildup and to remove the problemsassociated with the presence of the deposits by cleaning theturbine. The authors’ company has provided support to end userturbinesfor water washingof steamturbines(Watson, et al., 1995).

Theeffectiveness of thewater removal procedures mainly dependson the adherenceof thedeposits to thesubstrate.

A second route is to coat the surface with a material that hassuperior antifouling or antistick/corrosion characteristics. This inturnis beneficial totheturbinebladesby reducingthetendency forcontaminantstostick tothebladesandincreasetheeffectivenessof the water washing. Titanium nitride coatings with a chromiumundercoat (Cr-TiN) havealso been usedby steamturbineOEMs tocoat turbineblades.This coatingprovides corrosionprotectionandcan beused on all stages of a steam turbinerotor, however, theCr-TiN only provides limited antifoulant benefits. The authors’company has recently developed a proprietary coating, which isa corrosion resistant antifoulant coating designed for the laterstages of the turbine rotor (where the deposit buildup is mostsevere). Testing has shown that this coating provides signif icantimprovement in foulant release ability (Figure 55), excellentcorrosion protection (passes over 1000 hours of ASTM B117corrosion testing under a 5 percent salt solution), and erosionprotection (Figure 56), and, while having little effect on thefatigueproperties(Figure57).





Figure 55. Comparison of Foulant Release Performance of Bare Steel Against Proprietary Coating and Cr-TiN Coated Samples.

Figure 56. Compari son of Bare AISI 403 Stainl ess Steel Against Proprietary and Cr-TiN Coated Samples after 10 Hours of Modi fi ed ASTM G32 Testing.

Figure 57. Result s of R.R. Moore Fati gue Testi ng.

SUMMARY

An overview of materials and material related processes hasbeen presented for centrifugal compressors and steam turbines.Special attention has been given to address some of the problemsassociated with material selection for the various componentsandthe stepstaken to prevent and/or minimize reoccurrence.

What will the future bring to materials/or materials relatedprocesses? The application of composite materials for rotatingcomponents in compressors may be seen in the not so distantfuture. Theapplication of refinedexistingprocessestomanufacturecomponents to near net shapesuchas P/M net shape, HIP process,or metal rapidprototypingbasedonlaser microweldingof metallicpowders may also beseen.

ACKNOWLEDGEMENT

The authors are grateful for the support from the MaterialsEngineering Department at Elliott Company andrecognize ElliottCompany for permissionto publishthis paper.

REFERENCES

API RP 687, 2001, “Rotor Repair,” American PetroleumInstitute,Washington, D.C.

API Standard 617, 2002, “Axial and Centrifugal Compressors and

Expander-Compressors for Petroleum, Chemical and GasIndustry Services,” Seventh Edition, American PetroleumInstitute, Washington, D.C.

Aerospace Structural Metals Handbook: Volume 2, 1988, Departmentof Defense—Materials andCeramics Information Center.

Armor, A. F., Jaffee, R. I., andHottenstine, R. D., 1984, “AdvancedSupercritical Power Plants—TheEPRI DevelopmentProgram,”Proceedings of the Ameri can Power Conference, 46, p. 70.

Arlt, N., Gumpel, P., andSahni, P., 1988, “Uber EinigeEigenschaftenvon Nichtrostenden Nickelmartensitischen Chromstahlen,”Thyssen Edeist, Techn. Ber. 14, Band Heft 1, p. 71.

ASME Boil er & Pressure Vessel Code - Section VII I, 2007,AmericanSociety of Mechanical Engineers,NewYork, NewYork,pp. 66-67.

Atkins, M., 1980, Atlas of Continuous Cooling Transformation Diagrams for Engineeri ng Steels, Revised U.S. Edition,Copyright by theAmericanSociety for Metals, pp.168and188.

Batte, A. D. andMurphy, M. C., September 1981, “TheCorrosionFatigueof 12%Cr BladeSteels in LowPressureSteamTurbineEnvironments,” Corrosion Fatigue of Steam Turbine Blade Materials, Workshop Proceedings, Palo Alto, California.

Branch, G. D., et al., 1973, “High Temperature Bolts for SteamPower Plant,” International Conferenceon Creep and Fatiguein Elevated Temperature Applications, Sheffield andPhiladelphia, Institute of Mechanical Engineers, London,ConferencePublication 13, pp. 192.1-192.9.

Briggs, J. Z. and Parker, T. D., 1965, “The Super 12% Cr Steel,”Climax MolybdenumCo., NewYork, NewYork.

Bruscato, R. M., 1970, “Temper Embrittlement and CreepEmbrittlement of 2 Cr-1 Mo Shielded Metal-A rc WeldDeposits,” Welding Journal, 35, p. 148s.

Cameron, J. A., 1989, “Materials for Centrifugal Compressors—AProgressReport,”Proceedings of the Eighteenth Turbomachinery

Symposium, TurbomachineryL aboratory, TexasA&M University,CollegeStation,Texas, pp. 9-22.

Cameron, J. A., and Danowski, F. M., J r., 1973, “SomeMetallurgical Considerations in Centrifugal Compressors,”Proceedings of the Second Turbomachinery Symposium, Gas

TurbineLaboratories, TexasA&M University, CollegeStation, Texas, pp. 116-128.

Dowson, P., 1995, “Fracture Mechanics Methodology Applied toRotating Components of Steam Turbine and Centrifugal

Compressor Rotors,” Third International Charles Parson’sConference, Tyne, United Kingdom.

Dowson, P., 2002, “Development of 13Cr4Ni Material forBrazed/WeldedImpellersfor H2S andLowTemperatureService,”Supermartensitic Stainless Steels 2002, Brussels, Belgium.

Dowson, P., 2007, “Antifouling and CorrosionResistant Coatingsfor Cracked Gas Compressors and SteamTurbines,” AICHESpringNational Meeting, Houston,Texas.

Dowson, P., andWiegand, R., May 1996, “WeldedRotorRestoration— Technical Justification & Fit for Service Assurance,” SecondInternational EPRI ConferenceandVendor Exposition.




Dowson, P., Ross, S. L., andSchuster, C., 1991, “TheI nvestigationof Suitability of Abradable Seal Materials for Application inCentrifugal Compressorsand SteamTurbines,” Proceedings of the Twenti eth Turbomachinery Symposium, TurbomachineryLaboratory, Texas A&M University, College Station, Texas,pp. 77-90.

Dowson, P., Walker, M., andWatson, A., 2004, “Development of Abradable and RubTolerant Seal Materials for ApplicationinCentrifugal Compressorsand SteamTurbines,” Proceedings of the Thir ty-Thir d Turbomachinery Symposium, Turbomachinery

Laboratory, Texas A&M University, College Station, Texas,pp. 95-102.

Dowson, P., Wang, W., and Alija, A., 2005, “Remaining L ifeAssessment of Steam Turbine and Hot Gas ExpanderComponents,” Proceedings of theThir ty-Fourth Turbomachinery

Symposium, Turbomachinery Laboratory, Texas A&MUniversity, CollegeStation, Texas, pp. 77-92.

Engineered M aterial s H andbook: Composites, Volume 1, 1987,American Society of Metals, Materials Park, Ohio.

Everson, H., Orr, J., and Dulieu, D., 1988, “Low Alloy FerriticBolting Steels for SteamTurbineApplications: TheEvolutionof Durehete Steels,” in Advances in Materials Technology for Fossil Power Plants, R. Viswanathan and R. I. Jaffee, Editors,American Society for Metals, Metals Park, Ohio.

Fidler, F., 1971, “Metallurgical Considerations in WireWool TypeWear BearingPhenomena,” Wear, 17, pp. 1-20.

Gold, M. and Jaffee, R. I., 1984, “Materials for Advanced SteamCycles,” ASM Journal of Materials for Energy Systems, 6 (2),pp. 130-145.

Gross, J. H. and Stout, R. D., April 1957, “Ductility and EnergyRelations in Charpy Tests of Structural Steels,” Welding Research Supplement, 22, pp. 151s-159s.

Guinee, M. J. and Lamza, E. W., 1995, “Cost EffectiveMethods toMaintain Gas Production by the Reduction of Fouling inCentrifugal Compressors,” SPE30400, pp. 341-350.

Jones, G.T., 1972, Proceedings Instituteof Mechanical Engineers,186, p. 31-32.

Keller, H. F. and Cameron, J. A., 1974, “Laboratory Evaluation of Susceptibility to Sulf ide Cracking,” Carrier Corporation,NACE Paper #99.

Kohut, G. B. and McGuire, W. J., July 1968, “Sulfide StressCracking Causes Failure of Compressor Components inRefinery Service,” Materials Protection, 7, pp. 17-22.

LaFave, R. A., 1991, “SubmergedArc Weld Restoration of Steam Turbine Rotors Using Specialized Welding Techniques,”Proceedings of the Twentieth Turbomachinery Symposium,

TurbomachineryLaboratory, TexasA&M University, pp. 19-34.

LukensFinelineSteels, FormNo. 6754-85.,LukensSteel Company,Coatesville, Pennsylvania.

Meher-Homji, C. B., Focke, A. B., and Wooldridge, M. B., 1989,“Fouling of Axial Flow Compressors—Causes, Effects,

Detection, and Control,” Proceedings of the Eighteenth Turbomachinery Symposium, Turbomachinery Laboratory,

TexasA &M University, College Station, Texas, pp. 55-76.

Metals Handbook, Ninth Edition: Volume 4, 1981, AmericanSociety of Metals, Materials Park, Ohio, p. 35 and p. 43.

Moller, G. E., 1968, “Corrosion, Metallurgical and MechanicalExperiencesof PetroleumRefinery Compressors,”NACE TaskGroupT-8-1 Interim Report.

NACE Standard MR0103, 2005, “Materials Resistant to SulfideStressCracking in CorrosivePetroleumRefiningEnvironments,”NACE International, Houston, Texas.

NACE Standard MR0175, 2003, “Petroleum and Natural Gas

Industries—Materials for Usein H2S-Containing Environmentsin Oil and Gas,” NACE International, Houston, Texas.

NACE StandardRP0472, 2005, “Methods andControls to PreventIn-Service Environmental Cracking of Carbon SteelWeldments in Corrosive Petroleum Refining Environments,”NACE International, Houston,Texas.

Newhouse, D. L., et al., 1972, “Temper Embrittlement of AlloySteels,” ASTM STP 499, pp. 3-36.

Newhouse, D. L., J uly 1987, “Guide to 12Cr Steels for High-andIntermediate-Pressure Turbine Rotors for the AdvancedCoal-Fired Steam Plant,” Report CS-5277, Electric PowerResearch Institute, Palo Alto, California.

Schwant, R. C. andTimo, D. P., 1985, LifeAssessment of GeneralElectric Large Steam Turbine Rotors, Li fe A ssessment and

Improvement of Turbogenerator Rotors for Fossil Plants,Viswanathan, R., Editor, New York, New York: PergamonPress, p. 3-25-3-40.

Speidel, M. O., September 1981, “Corrosion-Fatigue of Steam Turbine BladeMaterials,” Corrosion Fatigueof SteamTurbineBladeMaterials, WorkshopProceedings, Palo Alto, California.

Treseder, R. S. and Swanson, T. M., 1968, “Factors in SulfideCorrosion Cracking of High Strength Steels,” Corrosion, 24,

pp. 31-37.

Viswanathan,R. andJ affee, R. I., October 1983, “Toughnessof Cr-Mo-V Steels for Steam Turbine Rotors,” ASME Journal of Engineering Material Techniques, 105, pp. 286-294.

Wang, W., Dowson, P., and Baha, A., 2003, “Developmentof Anti-

fouling and Corrosion Resistant Coatings for PetrochemicalCompressors,”Proceedings of t heThi rty-Second Turbomachinery

Symposium, Turbomachinery Laboratory, Texas A&MUniversity, CollegeStation, Texas, pp. 91-97.

Warren, D. and Beckman, G. W., October 1957, “SulphideCorrosion Cracking of High Strength Bolting Material,”Corrosion, 13, pp. 631t-646t.

Watanabe, J. and Murakami, Y., 1981, “Prevention of TemperEmbrittlement of CrMo Steel Vessels by the Use of Low SiForged Steels,” American Petroleum Institute, Chicago,Illinois, p. 216.

Watson, A. P., Carter, D. R., andAlleyne, C. D., 1995, “Cleaning Turbomachinery without Disassembly, Online and Offline,”Proceedings of the Twenty-Fourth Turbomachinery Symposium,

Turbomachinery Laboratory, Texas A&M University, CollegeStation, Texas, pp. 117-128.



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