72hvof coating and surface treatment for enhancing droplet

Wear 254 (2003) 652–667

HVOF coating and surface treatment for enhancing dropleterosion resistance of steam turbine blades

B.S. Mann∗, Vivek AryaSurface Coatings and Treatment Laboratory, BHEL, Corporate R&D Division, Vikasnagar, Hyderabad 500093, India

Received 14 May 2002; received in revised form 16 January 2003; accepted 16 January 2003

Abstract

This paper describes the water droplet erosion characteristics of high velocity oxygen fuel sprayed (HVOF) coated and laser hardened12Cr steels along with steels and titanium alloys used in steam turbine blades at two different energy fluxes. For droplet erosion study,round samples as per ASTM G73-98 were used. At low energy flux, the HVOF coated 12Cr steel performed much better than 12Cr and13Cr–4Ni steels. This is due to integrity of hard carbide particles in cobalt chrome matrix and its ability to absorb shocks due to highhardness of the carbide particles. During incubation as well as in the long run, laser hardened 12Cr steel performed exceptionally wellfollowed by 17Cr–4Ni ‘PH’ and heat-treated 12Cr steel. From the experimental study, it appears that ultimate and modified resilience ofmaterials play significant role to combat droplet erosion. Droplet erosion test results of all these materials and HVOF coating along withtheir properties and scanning electron micrographs are reported and discussed in this paper.© 2003 Elsevier Science B.V. All rights reserved.

Keywords:Steam turbine; Droplet erosion; HVOF coating; Laser hardening

1. Introduction

Water droplet erosion is a well-known phenomenon oc-curring on the moving blades operating at the low-pressure(LP) end of steam turbines. This is initiated by “small”, pri-mary droplet condensate in bulk of the supercooled steamin the flow, get separated on the blade surface and gener-ate secondary “large” droplets, which cause erosion[1–8].Relatively large drops 50–800�m diameter are consistentlyproduced and accelerated and strike on the convex side ofthe moving blades. The water drops greater than 200�mand a terminal velocity of more than 250 m/s are responsiblefor quick erosion[4,5,8,9]. Material loss from the leadingedges is the result of commutative damage from the impactof water droplets. The impact of large droplets 800�m at300 m/s generates high local forces. Higher velocity waterdroplet impact can produce serious erosion problems tohigh-speed military aircraft and missiles as well. Numerousmeans of combating droplet erosion are available in the liter-ature[6–27]. The damage produced by one or more of theseloading conditions on a material surface exposed to singleor multiple water drop impact is responsible for initiatingdamage and subsequent material removal. The evaluation of

∗ Corresponding author. Fax:+91-40-23776320.E-mail address:[email protected] (B.S. Mann).

damage produced in target material due to single water droploading cycle is a complex dynamic process, which involvesa number of closely phased actions. The damage mechanismis covered in reference[12–15]. Several important propertiesof materials such as material being cast, forged, rolled, an-nealed or in heat-treated condition including hard protectivelayers/coatings play an important role to combat impinge-ment erosion. Among all these, the hardness of mate-rial/protective layer or coating plays a significant role[27].

Corrosion-related failures of LP blades account forone-third of the total number of blade failures in thelow-pressure turbine. These are the main cause of forcedoutages of the steam turbine[16]. Corrosion fatigue alongwith pitting is considered to be the most common singlefactor. The formation of pits due to droplet erosion at acritical location on a turbine blade can act to aggravatecracking in two ways: as a stress concentration causing theblade to experience high stresses (both static and cyclic)and as a stagnant corrosion cell where the environment atthe base of the pit can change concentration as the corrosionproceeds. The latter situation can lead to a lower pH, moreacidic environment known to be deleterious to the corrosionfatigue life of conventional blading alloys. The corrosionerosion characteristics in 3.5% NaCl solution at 23◦C oflaser hardened steel were studied. This is confirmed bythe pitting potential measurements of the specimens. The

0043-1648/03/$ – see front matter © 2003 Elsevier Science B.V. All rights reserved.doi:10.1016/S0043-1648(03)00253-9

B.S. Mann, V. Arya / Wear 254 (2003) 652–667 653

Table 1Room temperature tensile properties of LP turbine blading alloys[16]

Materials Tensile strength(MPa)

Yield strength(MPa)

Elongation (%) RA (%) Fracture toughness(MPa m2)

12% Cr steel (UNS S403) 759–786 645–672 20.0–20.5 63.3–67.1 18017Cr–4Ni PH steel (UNS S174) 952–972 656–712 19.6–20.7 67.4–68.5 170Titanium alloy Ti–6Al–4V 1095 1013 12.0 34.5 67

Table 220 kHz 109 cycle fatigue strength in corrosive media[16]

Materials Fatigue strength(MPa) (a)

Fatigue strength(MPa) (b)

12 Cr steel 390 10017Cr–4Ni PH steel 430 180Ti–6Al–4V 470 290

Test conditions: (a) triple deionised water, 30◦C, pH 7, <7 ppb O2; (b)22% NaCl, 80◦C, pH 4,<20 ppb O2.

improvement in pitting corrosion resistance resulted fromthe dissolution or refinement of carbide particles and thepresence of retained austenite, as evidenced by the fact thatthe pitting potential increased linearly with the amount ofretained austenite.

Typical blade alloys considered for LP turbine are listedin Table 1. The LP turbine blades are martensitic stainlesssteels. The austenitic and ferritic steels are not consideredsuitable for LP turbine blades on the grounds of suscepti-bility to cracking and lack of strength. For design purposes,the yield strength is the most important factor for assessingstatic strength. Ductility is also an important criterion alongwith static strength to accommodate localized stress peaksand stress concentrations. When wet steam becomes impure,the fatigue strengths and residual stress state of blade materi-als are affected very largely[29]. Under resonant conditions,blade deflections and consequent stresses become very highso that it may fail due to mechanical fatigue. Therefore, thealloys should have adequate fatigue strength in the operat-ing environment. The fatigue strength of various LP bladealloys is shown inTables 2 and 3. To combat these damagesthe surface modification of the components is very effectiveand attractive since it does not affect the bulk properties ofthe substrate.

Titanium blades are also in use since long. The dropleterosion studies on Ti–6Al–4V alloy revealed that the ero-sion resistance of this alloy is comparable to the erosionresistance of 12Cr steel, however slightly inferior to flame

Table 3Fatigue Strength, 107 cycles at 100 Hz[16]

Materials Fatigue strength (MPa)

12Cr steel 10017Cr–4Ni PH steel 350Ti–6Al–4V 400

22% NaCl, 80◦C, pH 4,<20 ppb O2.

hardened 12Cr steel in the velocities range 300–500 m/s anddroplet size 0.10–0.80 mm[30]. Some information on cav-itation erosion resistance of laser hardened steel is avail-able[28,29]. However, not much information on the dropleterosion resistance of laser hardening of steel is available.Droplet erosion studies on laser nitrided Ti–6Al–4V alloyhave also been reported[30,31]. An improvement of the or-der of 350–400% was observed on laser nitrided Ti–6Al–4Valloy.

Tests conducted by Brown and Westinghouse have shownthat the steam moisture content and peripheral velocity lim-itations imposed on 12Cr steel (type 403 steel) blades canbe considerably extended with Ti–6Al–4V blades as well.A slight amount of erosion of titanium as well as 12Cr steelblades after an operation of 4 years was observed whereasstellited steel blades were unaffected. Stelliting is generallynot used because of thermal expansion mismatch betweenthe two alloys. In addition to steady loads from rotation andsteam flow, blades are subjected to vibratory excitations aswell. Vibratory stresses result from the vibration of untunedmodes, which are excited by circumferential non-uniformityin the steam flow, which are difficult to be assessed.

In this paper, the results on our recent experimental find-ings on droplet erosion resistance of high velocity oxygenfuel sprayed (HVOF) coating and laser hardened 12Cr steelalong with steam turbine blading materials such as tita-nium Ti–6Al–4V alloy, 12Cr steel (AISI 403 and 410 type),17Cr–4Ni PH steels in as received (‘AS’) and in hardenedcondition (‘HT’) and 13Cr–4Ni steel in as received (‘AS’)condition are reported.

2. Erosion prediction model

For erosion prediction Haymann proposed the followingequation[7]

UeM = Ua1

Na

(V0

2550

)5

This equation gives maximum instantaneous value of thematerial volume loss per unit area per unit time. The equationis based upon the experimental data. The droplet structurewas assumed homogenous and the angle of attack was about90◦. This equation could not be applied directly on turbineblade, where every blade element is being hit by a relativelybroad spectrum of droplet sizes. Moreover, every droplet

654 B.S. Mann, V. Arya / Wear 254 (2003) 652–667

group defined by a given droplet radiusr∗ hits the surface ofthe blade element at a different angle of attack. This equationis further modified by Krzyzanowski and Szperngiel[8],which takes care of all these points.

UeM(η) =r∗(η)∑

r∗=r∗,min

sinβb1(η)δM∗(r∗, η)

ρ dη�R1

(w∗N(r∗, η)

2550

)2

whereUeM(η) represents material volume loss per unit areaper unit time;Ua or δM∗(r∗, η)/ρ dη�R1 represents massflux; V0 orw∗N(r∗, η) represents velocity component;βb1(η)

represents impingement angle; (r∗, η) represents dropletsize;Na represents normalization with respect to 18-8 stain-less steel.

In the present study, the theoretical prediction for erosiondamage estimation has not been done. However, all the pa-rameters such as energy flux, mass flux, impingement angle,droplet size and velocity were considered while doing theexperiment. Their values are given inTable 5.

3. High velocity oxygen fuel (HVOF) coatingsand laser hardening

3.1. HVOF coatings

In recent years, high velocity oxygen fuel coating hasbeen considered an asset to the family of thermal spray pro-cesses, especially for thermal spraying of materials with amelting point below 3000◦ K. It has proven successful, sinceit shows advantages in density, bond strength and makingit attractive for many wear as well as corrosion resistanceapplications[32–37]. Its high coating quality results fromthe use of a hot combustion-driven compared to chemi-cally/electrochemically formed coatings.

Tungsten carbide (WC) powders are widely used in thehigh velocity oxygen fuel spraying system[33]. These areused to produce dense, high hardness and excellent wear re-sistance coatings generally to combat erosion and corrosionoccurring in the industry. In applications, where abrasive orerosive wear resistance is of primary importance, WC–Cowith and without nickel or chrome is used. WC–Co–Cr pow-ders are preferred when there is a high demand for corro-sion resistance. The erosion resistance depends upon oxidesand pores, and the phase transformation occurring duringspraying.

High velocity oxygen fuel sprayed coatings are commonlyapplied by HP/HVOF JP-5000, DS-100, Met jet II, OSU,Diamond jet and Praxair 2000 HVAF systems. These sys-tems are based on liquid fuel and oxygen, gaseous fuel andoxygen, and liquid fuel and air. It has been reported thatcoating microstructure, hardness and composition were themajor determinants in erosive wear. The matrix corrosionalso influences the erosive wear.

Surface preparation before HVOF coating is a very im-portant step in thermal spraying. This is because the ad-

hesion of coatings is directly related to the roughness ofthe surface and it is controlled by the type of grit blastingmachine, blasting pressure, angle, distance, time and gritblasting nozzle (orifice size)[38]. The round specimensof size ∅ 12.7 × 40 mm long were degreased by carbontetrachloride vapour degreasing technique. These weregrit blasted using suction type grit blasting machine usingalumina grit of size 24 mesh. The coatings were sprayedusing Met Jet II HP/HVOF system. This system is basedupon liquid fuel kerosene and oxygen. The parametersadopted while surface preparation. HVOF spraying param-eters are within the range, which are generally adopted forapplying an HVOF coating[39]. In brief these are givenbelow.

Combustion pressure 0.72 MPaSpray distance 380 mmSpray angle 90◦Powder feed rate 70 g/minFuel flow rate 24.75 l/hOxygen flow rate 950 l/mCarrier gas flow rate 45 l/mBarrel length 150 mmSpraying powder WC–10Co–4Cr, Praxair WC 636

HVOF coating on all the round samples size∅ 12.7 ×40 mm long was given in automatic mode. The coating thick-ness in the range of 250± 30�m was maintained on all thesamples.

3.2. Laser hardening

Laser hardening studies were limited to a narrow powerdensity in the range of 1740–2400 W/(cm2/s). After con-ducting all the tests and confirming all the above findings,a laser power of 2120 W/(cm2/s) was selected for exper-imentation. These studies were based upon Kwok et al.detailed experimentation on laser hardening of 12Cr steel[28].

4. Water droplet erosion testing of different coatings

As per ASTM G73, water impingement erosion testfacility has been fabricated in house. After establish-ing the accuracy of results similar to those reported inASTM G73, this facility has been used for testing ofmaterials/coatings. The test facility consists of a cham-ber of diameter 700 mm, and a round disc on which thetest samples are fixed on the periphery. The dimensionaldetails of the test rig are given inFig. 1. The disc is de-signed for rotating in a wide range of speeds from 3000to 6000 rpm. Two water jets impinge on the cylindricaltest samples and cause impingement erosion. The cylin-drical specimens were selected because the impingement


Fig. 1. Cross-sectional view of droplet erosion test rig.

erosion on actual LP steam turbine blades occurs at theleading edge. Two different energy fluxes, 28.8 × 106

and 32.64 × 106 J/m2 s were selected to study the ero-sion damages occurring in HVOF coated and Laser treatedsteels.

A precision balance to an accuracy of 0.1 mg was usedfor measurement of weight loss occurring after a certain testduration. The test duration was selected in such a way thatsteady state impingement erosion occurred. The accuracy ofthe test results has been confirmed using a reference 12Crsteel. The deviation and accuracy lie within specified accu-racy data already available from different laboratories[13].The results have been plotted as commutative erosion–timecurve on a co-ordinate of mean depth of erosion versustime. The depth of erosion is calculated from the weightloss divided by the density of the coatings and the materials.Table 4gives the materials tested for droplet erosion. Testconditions are given inTable 5.

Table 4Various materials used for droplet erosion testing

Materials Composition (wt.%)

12Cr ‘AS’ 0.10C, 12Cr, 0.6Si, 0.70Mn Balance Fe12Cr (ST ‘AS’) 0.20C, 12Cr, 0.5Si, 0.5Mn, 0.5Ni Balance Fe13Cr–4Ni 0.058C, 12Cr, 0.5Si, 0.5Mo, 3.85Ni Balance Fe17Cr–4Ni PH ‘AS’ 0.06C, 15.67Cr, 0.27Si, 0.64Mn,

4.25Ni, 3.6Cu, 0.19NbBalance Fe

Ti–6A1–4V 6Al, 4V Balance Ti

Table 5Experimental test conditions

Conditions Test I Test II

Water jet size 3.0× 10−3 m 3.77× 10−3 mWater mass flux 2.65 m/s 3 m/sWater energy flux 28.8× 106 J/m2 s 32.64× 106 J/m2 sRelative water velocity 147.4 m/s 147.7 m/sTest sample size ∅ 12.70× 40 mm ∅ 12.70× 40 mmWater droplet size 100–300�m 100–300�mNumber of specimens used 10 10Test duration 5.49× 106 N 2.745× 106 NJet distance 100 mm 100 mmAngle of impingement 0–90◦ 0–90◦Impact frequency 78 cycles/s 78 cycles/sExperimental accuracy ±21.25% ±17.5%

5. Results and discussions

5.1. Scanning electron micrographs

Scanning electron microstructures of eroded samples weretaken at low and high magnifications. These are given inFigs. 2–13. From the micrographs, it is observed that thatthe damage mechanism of 12Cr ‘AS’, 12Cr (ST ‘AS’), and13Cr–4Ni ‘AS’ steels is identical and all these materials arefailing due to formation of pits due to cleavage. The cleavagedamage in 13Cr–4Ni ‘AS’ steel is slightly less comparedto the other two steels. The HVOF coating has failed inbrittle mode in laminates (Figs. 12 and 13). This is because


Fig. 2. SEM of eroded 12Cr (ST ‘AS’) steel (30×).

the HVOF coating is deposited layer by layer. The grainmorphology of 17Cr–4Ni PH ‘AS’, 17Cr–4Ni PH ‘HT’ and12Cr (ST ‘HT’) is identical and similar. 17Cr–4Ni PH ‘HT’is the finest among all these followed by 17Cr–4Ni PH ‘AS’,and 12Cr (ST ‘HT’). Because of this 17Cr–4Ni PH ‘HT’ hasperformed much better than other steels. 17Cr–4Ni PH series(17Cr–4Ni PH ‘AS’ and 17Cr–4Ni PH ‘HT’) are erodingin the ductile mode. No oxide deposits were seen on, 12Cr(ST ‘HT’), 12Cr ‘LH’, 17Cr–4Ni PH ‘AS’, 17Cr–4Ni PH‘HT’ and Ti–6Al–4V. Relevant micrographs confirm thatcorrosion does not contribute to erosion of these materials.SEM of Ti–6Al–4V shows that the grains are fine, so itsperformance is also similar to 17Cr–4Ni PH series.

Fig. 3. SEM of eroded 12Cr (ST ‘AS’) steel showing coarse structure (1.5 K).

From the micrographs, it is seen that the mode of mate-rial removal in 17Cr–4Ni PH series and 12Cr (ST ‘HT’) isacross the grain and in case of Ti–6Al–4V, deep cavities areobserved in longitudinal direction (across the flow) whichconfirm that the variation of mechanical properties in longi-tudinal directions compared to transverse direction.

The micrographs of 12Cr ‘LH’ are similar to 12Cr (ST‘HT’). Retention of austenitic in laser processing does notallow the material to erode easily[28]. Due to this, 12Cr‘LH’ is performed excellent compared to all other materialsat both the energy flux levels.

At high energy flux, 13Cr–4Ni ‘AS’ steel has performedmuch better than 12Cr (ST ‘AS’) and 12Cr ‘AS’. This is


Fig. 4. SEM of eroded 12Cr (ST ‘HT’) steel (30×).

also because of a finer microstructure compared to othersteels. After long exposure, deep microtunnels confirmingmicrojetting effects similar to cavitation erosion mechanismare observed in these steels, whereas in the 17Cr–4Ni PH,12Cr ‘LH’ and 12Cr (ST ‘HT’) steels these deep microtun-nels confirming microjetting effects are not observed. Thesemicrojetting effects in these steels may appear after longtesting.

5.2. Properties of materials and coatings

Tables 6 and 7give the mechanical properties of all thematerials and coatings. All these properties were measuredusing tensile and hardness testing machines. From the me-

Fig. 5. SEM of eroded 12Cr (ST ‘HT’) steel showing very fine grain structure (1.5 K).

chanical properties the best property which has given anexcellent correlation with erosion resistance is modified re-silience, a product of ultimate tensile strength (UTS) andhardness. 12Cr (ST ‘HT’) followed by 17Cr–4Ni PH seriesfollows the trend of droplet erosion resistance. Other prop-erties, such as toughness, yield strength do not play muchrole. The tensile strength of laser hardened 12 Cr, and HVOFcoating could not be measured although some data on tough-ness on HVOF coating is available[32].

5.3. Microhardness results

The microhardness of the coated steels was mea-sured using Leitz’s Micro Hardness Tester by applying a


Fig. 6. SEM of eroded 12Cr ‘LH’ steel showing damage due to individual droplets (75×).

load of 2.942 N. The microhardness values are given inTable 6.

5.4. Surface finish

After an operation of 7 h, the surface finish of erodedsamples was measured by a perthometer. These values aregiven in Figs. 14–19. From figures it is seen that the 12Cr‘AS’, 12Cr (ST ‘AS’) and 13Cr–4Ni ‘AS’ samples havebecome very rough. EDX analysis shows that the calciumand sodium salt deposits are more in rough samples. Dueto this, the phenomenon of erosion and corrosion has beenaccelerated in these steels.

Fig. 7. SEM of eroded 12Cr ‘LH’ steel showing very fine grain structure similar to ST ‘HT’ (1.5 K).

5.5. X-ray diffraction test results

X-ray analysis of WC636 powder reveals tungsten car-bide as a main phase and small percentage of Co6W6C alongwith Co and Cr as binder. After HVOF coating, WC was par-tially converted into W2C (only small percentage 4–5%) andCo6W6C has been recorded in traces. Cobalt and chromiumhave converted into amorphous phases.

5.6. Droplet erosion test results

The erosion test results of different stainless steels andtitanium alloy along with HVOF coating are given in


Fig. 8. SEM of eroded 17Cr–4Ni PH ‘HT’ steel showing very fine grain structure (1.5 K).

Figs. 20–22. It is seen from the figures that excellent per-formance is given by laser hardened 12Cr steel followed by17Cr–4Ni PH series and 12Cr (ST ‘HT’). Kwok et al. hasreported that excellent resistance to cavitation erosion incase of laser treated stainless steel and AISI 420 steel[28].This was obtained at a power density of 2266 W/(cm2/s).The austenitic phase has an excellent characteristic of ab-sorbing impact energy. Conversion of austenitic phase intomartensitic phase has induced compressive stresses on thesurface, which are beneficial to erosion. Evidence of trans-formation of austenitic phase into martensitic phase aftercavitation was also observed. Potentiodynamic polarisation

Fig. 9. SEM of eroded 17Cr–4Ni PH ‘AS’ steel showing fine grain structure (1.5 K).

studies show that the pitting corrosion is low in laser treatedsteel because laser hardening has resulted in complete dis-solution of carbides. The corrosion studies carried out ondifferent volume fractions of retained austenitic also provedthat corrosion rate decreased linearly with increased volumefraction of retained austenitic phase during laser hardening.The sizes of the pits on 12Cr (ST ‘AS’) steel were mea-sured under the microscope. The sizes are in the range of30–100�m similar to the pits that generally occur duringactual droplet erosion of the steam turbine blades. Thesepits corresponds to droplet size in the range of 100–300�m.Gram and Sturm[29] have reported that laser hardening


Fig. 10. SEM of eroded Ti–6Al–4V alloy (30×).

introduces high tensile stresses of the order of 200 MPadue to rapid heating and cooling. These stresses are detri-mental in accelerating stress corrosion-related fatiguesdamages.

It can be seen fromFig. 22 that HVOF sprayed WC636steel has performed much better at low energy flux comparedto higher energy flux whereas 12Cr ‘LH’, 12Cr (ST ‘HT’),and 17Cr–4Ni PH series have performed well at all the en-ergy fluxes including incubation period. For all the materials,a drastic reduction in volume loss is observed when energyflux is reduced from 32.64× 106 to 28.8 × 106 J/m2 s butthis reduction is very significant in case of HVOF coating.

Fig. 11. SEM of eroded Ti–6Al–4V alloy showing fine grain structure (1.5 K).

Incubation period and ranking of the coatings and materialsare given inTables 8–10.

In actual steam turbine as reported by Krzyzanowski andSzperngiel[8], the energy flux values are much lower so theperformance of HVOF coating will improve further signif-icantly. The HVOF coating requires field evaluation as thelaboratory testing for long duration to the steam turbinesoperating conditions may not be practical and possible.

At low energy flux, the superior performance of HVOFcoating is due to the hardness of tungsten carbide particles(1800 HV). These are well embedded in the matrix and donot allow wearing out of matrix as these come directly in


Fig. 12. SEM of eroded HVOF coating showing layer by layer damage (30×).

Fig. 13. SEM of eroded HVOF coating (1.5 K).

Fig. 14. Surface roughness of 12Cr (ST ‘AS’) steel before erosion.


Table 6Mechanical properties of different coatings and materials

Materials/coatedmaterials

Yield strength(N/mm2)

Ultimate tensile(N/mm2)

Hardness(HV)

Elongation(%)

Ultimate resilience(J cm−3)

Strain energy(J cm−3)

Impactstrength (J)

12Cr ‘AS’ 464.79 720.45 200–210 26.98 1.23 157.3 15312Cr (ST ‘AS’) 721.0 876.0 300–350 23.17 1.83 164.3 93.012Cr (ST ‘HT’) 1134.0 1562.7 450–500 15.6 5.78 197.27 46.213/4 ‘AS’ 813.2 892.20 300–310 14.8 1.90 112.1 78.017Cr–4Ni PH ‘AS’ 863.49 1224.28 365–380 13.44 3.577 139.68 112.017Cr–4Ni PH ‘HT’ 1305.09 1448.68 450–460 13.04 5.0 160.3 38Ti–6Al–4V 850 874 330–350 13 3.17 96.0 6512Cr ‘HVOF’ – – 1090–1226 – – – –12Cr ‘LH’ – – 550–650 – – – –

Ultimate resilience: UTS2/2E; UTS: ultimate tensile strength;E: Young’s Modulus. The 12Cr ‘AS’ in annealed condition. The 12Cr (ST ‘AS’): actualsteam turbine blade material in a forged condition and later on stress relieved for 4 h at 250◦C. The 12Cr (ST ‘HT’): actual steam turbine blade materialin forged condition and heat-treated at 950◦C for 1 h followed by water quenching and later on stress relieved for 4 h at 250◦C. The 13/4 ‘AS’ in ascast condition and later on stress relieved for 4 h at 250◦C. The 17Cr–4Ni PH ‘AS’ in cast condition and later on stress relieved for 4 h at 250◦C.The 17Cr–4Ni PH ‘HT’ in as cast condition and later on aged for 3 h at 490◦C. Ti–6Al–4V alloy in forged condition and later on heat-treated for 1 hat 950◦C followed by water quenching and aged for 6 h at 535◦C. 12Cr ‘HVOF’: HVOF coated 12Cr steel having fracture toughness in the range of4.5–5 MPa m2 [32]. 12Cr ‘LH’: laser hardened 12Cr steel. Microhardness of 12Cr HVOF was taken at 300 g load.

Fig. 15. Surface roughness of 12Cr (ST ‘AS’) steel after erosion.

contact with the water droplets due to higher volume frac-tion. It is well known that physical properties of materialssuch as ultimate tensile strength, modified resilience, bind-ing energy and crystal structure play a crucial role in de-termining the erosion. Feller and Kharrazi[22] have shownthat the higher the yield strength and crystal binding energyof a material, the longer the incubation period. However, a

Fig. 16. Surface roughness of 12Cr (ST ‘HT’) steel before erosion.

prerequisite is the integrity of the coatings together with ahigh degree of bonding to the substrate. Diffused coatingsproduced by chemical vapour deposition are the techniquesof achieving such composite structures as reported by Frees[23]. The impingement erosion resistance of HVOF coatingscan be explained on the basis that microjets formed by di-viding bigger jets into smaller jets cannot penetrate the hard


Fig. 17. Surface roughness of 12Cr (ST ‘HT’) steel after erosion.

Fig. 18. Surface roughness of HVOF coating before erosion.

Fig. 19. Surface roughness of HVOF coating after erosion.

tungsten carbide particles easily. The repetitive high in-tensity of the striking jet weakens the matrix and is thenremoved easily by microjets. The significant improvementof HVOF coated substrates at low energy flux indicatesthat the hardness of coatings has a crucial role to play intheir performance. Catastrophic damage at higher energylevels may follow the trends observed for hard and brittlematerials in cavitation erosion such as tungsten carbideand Haynes alloy 6B[12]. From Figs. 20 and 22, it isseen that the performance of HVOF coating has improvedsignificantly by reducing the energy flux and it will im-prove significantly further at lower energy fluxes. Kurodaet al. have reported that compressive stresses ranging from

Table 7Experimentally determined properties of different materials

Materials Hardness(HV)

Modifiedresilience (HV)

Impactstrength (J)

12Cr ‘AS’ 200–210 0.369 15312Cr (ST ‘AS’) 300–350 0.712 93.012Cr (ST ‘HT’) 450–500 1.86 46.213/4 ‘AS’ 300–310 0.68 78.017Cr–4Ni PH ‘AS’ 365–380 1.13 112.017Cr–4Ni PH ‘HT’ 450–460 1.68 38Ti–6Al–4V 330–350 1.29 65

modified resilience= UTS(substrate) × hardness(substrate)

2E.


Fig. 20. Volume loss of different materials and coatings at an energy flux of 28.8×106 J/m2 s.

70 to 420 MPa were found to be generated during HVOFspraying due to peening effect of spraying particles insemimolten state at high velocity for different materialssuch as 316 stainless steel, Haste alloy C, WC–12% Coon stainless steel substrate. High compressive stresses ofthe order of 200 MPa in HVOF coating and up to 50�mdeep inside the substrate are common[40]. These com-

Fig. 21. Volume loss of different materials and coatings at an energy flux of 32.64×106 J/m2 s.

pressive stresses generated were found to increased withkinetic energy of the spraying particles and are highlybeneficial in reducing stress corrosion-related fatiguesdamages.

From Tables 6 and 7andFigs. 20–22, it is seen that thematerial having higher ultimate and modified resilience hasperformed extremely well in droplet erosion. This property


Fig. 22. Droplet erosion test results of different materials/coatings at 32.64 and 28.8×106 J/m2 s.

Table 8Incubation period of different coatings/materials at different energy fluxes

Coatings/materials Energy flux of 28.8× 106 J/m2 s (h)

Energy flux of 32.64× 106 J/m2 s (h)

12Cr ‘AS’ 1 <113Cr–4Ni ‘As’ 2 3HVOF coated 12Cr 3 212Cr (ST ‘AS’) 5 2Ti–6Al–4V 6 312Cr (ST ‘HT’) 7 317Cr–4Ni PH ‘AS’ 8 417Cr–4Ni PH ‘HT’ 8.5 412Cr ‘LH’ >20 8

Table 9Volume loss and ranking of different coatings/materials after 2.745×106 Ncorresponding to an energy flux of 32.64× 106 J/m2 s

Coatings/materials Ranking Volume loss (mm3)

12Cr ‘As’ 9 >43.512Cr (ST ‘AS’) 8 26.42HVOF coated 12Cr 7 15.1713Cr–4Ni ‘As’ 6 13.8Ti–6Al–4V 5 6.0717Cr–4Ni PH ‘AS’ 4 4.112Cr (ST ‘HT’) 3 2.1517Cr–4Ni PH ‘HT’ 2 1.8912Cr ‘LH’ 1 0.82

Table 10Volume loss and ranking of different coatings/materials after 2.745×106 Ncorresponding to an energy flux of 28.8 × 106 J/m2 s

Coatings/materials Ranking Volume loss (mm3)

12Cr ‘As’ 9 17.813Cr–4Ni ‘As’ 8 3.02ST ‘As’ 7 1.68HVOF coated 12Cr 6 1.14Ti–6Al–4V 5 0.9312Cr (ST ‘HT’) 4 0.517Cr–4Ni PH ‘AS’ 3 0.3617Cr–4Ni PH ‘HT’ 2 0.1412Cr ‘LH’ 1 0

may be right choice to select a blade material to combatdroplet erosion.

6. Conclusions

1. The HVOF coating has failed in brittle mode. It has comeout layer by layer. This is because the HVOF coatinghas been deposited layer by layer. By reducing dropletenergy flux the performance of HVOF coating has im-proved significantly. For lower energy flux values similarto those occur in steam turbine blades, the HVOF coatingmay be an appropriate solution for droplet erosion. This


technique has an advantage of introducing compressivestresses within coatings as well as in the base material,which are beneficial to improve fatigue-related corrosiondamages.

2. Laser-hardened 12Cr steel has shown excellent perfor-mance. This is due to the retention of higher austeniticphase and complete dissolution of carbides as reported byKwok et al. The austenitic phase has an excellent char-acteristic of absorbing water impact shocks. Later on,it is converted into a martensitic phase. Conversion ofaustenitic phase into martensitic phase has induced com-pressive stresses on the surface, which are beneficial toovercome impingement erosion.

3. The grain morphology of 17Cr–4Ni PH ‘AS’, 17Cr–4NiPH ‘HT’ and 12Cr (ST ‘HT’) is identical and similar.17Cr–4Ni PH ‘HT’ and 12Cr ‘LH’ are fine among allthese followed by 17Cr–4Ni PH ‘AS’, and 12Cr (ST‘HT’). Because of this, 17Cr–4Ni PH ‘HT’ and 12Cr‘LH’ have performed much better than other steels.17Cr–4Ni PH series (17Cr–4Ni PH ‘AS’ and 17Cr–4NiPH ‘HT’) erode in the ductile mode. No oxide depositswere seen on Ti–6Al–4V, 12Cr (ST ‘HT’), 17Cr–4Ni PH‘AS’, 17Cr–4Ni PH ‘HT’ and 12Cr ‘LH’. This confirmsthat corrosion does not contribute to erosion for thesematerials.

4. SEM of Ti–6Al–4V shows that the grains are fine, soits performance is also similar to 17Cr–4Ni PH series.From the micrographs, it is seen that the mode of mate-rial removal in 17Cr–4Ni PH series and 12Cr (ST ‘HT’)is across the grain whereas in case of Ti–6Al–4V, deepcrates are observed in the longitudinal direction (acrossthe flow) which confirms that there is a variation of me-chanical properties of this materials in the longitudinaldirection compared to transverse direction.

5. 13Cr–4Ni steel has performed much better than 12Cr (ST‘AS’) and 12Cr ‘AS’ at higher energy flux. This is be-cause of finer microstructure compared to other steels.After long exposure, deep microtunnels confirming mi-crojetting effects mechanism are observed in all thesesteels, whereas in 17Cr–4Ni PH series, 12Cr ‘LH’ and12Cr (ST ‘HT’) these deep microtunnels confirming mi-crojetting effects are not observed.

6. 12Cr (ST ‘AS’) steel has performed much better than12Cr ‘AS’ at both the energy flux values. This is due toits higher values of ultimate and modified resilience. Thiscriterion is also valid for 17Cr–4Ni PH series and 12Cr(ST ‘HT’) and Ti–6Al–4V.

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