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Mo and W as alloying elements in Co-based alloys—their effects on erosion–corrosion resistance
Wear 259 (2005) 219–229
Mo and W as alloying elements in Co-based alloys—their effects on erosion–corrosion resistance Ugur Malayoglu, Anne Neville ∗ Corrosion and Surface Engineering Research Group, School of Mechanical Engineering, University of Leeds, Leeds LS2 9JT, UK Received 27 July 2004; received in revised form 27 January 2005; accepted 2 February 2005 Available online 10 May 2005
Abstract In this paper results from erosion–corrosion tests performed under liquid–solid erosion conditions in 3.5% NaCl liquid medium are reported. The focus of the paper is to compare the behaviour of Stellite alloy 706 which contains 5 wt.% molybdenum with Stellite alloy 6 which contains 4.8 wt.% tungsten, both in cast and Hot Isostatically Pressed (HIPed) form, in terms of their electrochemical corrosion characteristics, their resistance to mechanical degradation and relationship between microstructure and degradation mechanisms. It has been shown that both cast and HIPed Stellite 706 possess better erosion and erosion–corrosion resistance than Stellite 6 counterparts under a wide range of conditions. The significant role of the microstructure and specifically the type of carbides in affecting the erosion–corrosion performance of the alloys is demonstrated and discussed. Also it has been shown that in these multiphase alloys there is no direct relationship between the hardness and erosion–corrosion resistance. © 2005 Elsevier B.V. All rights reserved. Keywords: Co-based alloys; Liquid–solid slurries; Erosion; Corrosion
1. Introduction In the chemical, petrochemical and pump industries, machine parts often work in conditions where erosion and corrosion processes, acting together, are the main failure mechanisms. The deleterious synergistic effect of the erosion removes either corrosion product or the passive layer that is protecting the underlying surface. When a passive layer is removed, the time taken for it to repassivate is an important consideration in the assessment of wear rates. The rate of repassivation determines the amount of charge that can transfer when the surface is activated by an erosion event (e.g. impact of sand). The rate of repassivation in relation to the frequency of impact is an important consideration in erosion–corrosion. On stainless steels [1] it was shown that higher amounts of key elements (chromium, molybdenum) can lead a faster repassivation during erosion impacts.
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Corresponding author. Tel.: +44 113 343 6812; fax: +44 113 242 4611. E-mail address: [email protected] (A. Neville).
0043-1648/$ – see front matter © 2005 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2005.02.038
Cobalt-based alloys have enjoyed extensive use in wearrelated engineering applications for well over 50 years because of their inherent high-strength, corrosion resistance and ability to retain hardness at elevated temperatures [2]. In recent years a concentrated effort has been made to understand the deformation characteristics of cobalt-based alloys exposed to erosion–corrosion environments in order to optimize those factors contributing to their erosion resistance [3–5]. Alloying cobalt with chromium and various quantities of carbon, tungsten and molybdenum produces a family of alloys which can have excellent resistance to corrosion and/or erosion. Understanding how microstructural changes, as a result of alloying, affect corrosion and erosion resistance is critical to optimising the alloy for a particular purpose. In cobalt-based alloys, the key element chromium is added in the range of 20–30 wt.% to improve corrosion and impart some measure of solid-solution strengthening. Where carbide precipitation strengthening is a desirable feature, chromium also plays a strong role through the formation of a series of varying chromium–carbon ratio carbides such as M7 C3 and M23 C6 . Alloying elements like tungsten, molybdenum and
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tantalum are added to cobalt for solid solution strengthening. If these metals are added in excess of their solubility, formation of carbides like MC and M6 C is likely to occur. Shin et al. [6] investigated the effect of molybdenum on the microstructure and wear resistance properties of Stellite 6 hardfacing alloy. They showed that with an increase in molybdenum content, the M23 C6 and M6 C type carbides were formed instead of chromium-rich M7 C3 . They concluded that this microstructural change was responsible for the improvement of the mechanical properties such as hardness and wear resistance of molybdenum-modified Stellite 6 hardfacing alloy. Most cobalt-based alloys possess outstanding cavitation resistance compared to stainless steels which has been shown to be independent of the carbon content (hence hardness), and has been attributed by Crook [7] to crystallographic transformation, under stress, from the face centred cubic (fcc) to hexagonal close packed (hcp) structure by twinning. Heathcock and Ball [8] studied the cavitation resistance of a number of Stellite alloys, cemented carbides and surface-treated alloy steels and showed that in Stellite alloys the cobalt-rich solid solution, incorporating elements such as chromium, tungsten and molybdenum is highly resistant to erosion due to a rapid increase in the work-hardening rate and the strain to fracture which are caused by deformation twinning. Lee et al. [9] compared the liquid impact erosion resistance of 12 Cr steel with a Vickers hardness of 380 kg/mm2 (∼39 MPa) and Stellite 6B with a hardness value of 420 kg/mm2 (∼43 MPa). The liquid impact erosion resistance of Stellite 6B was at least six times greater than that of 12 Cr steel, implying that hardness is not the governing factor for liquid erosion. Stellite 6B also showed very different behaviour in liquid impact erosion in comparison with 12 Cr steel. They concluded that the superior erosion resistance of Stellite 6B results from the cobalt matrix whose deformation appeared mostly as mechanical twins and the material removal was more dominant in the hard carbide precipitates than in the ductile cobalt matrix. Wong-Kian [10] showed that Stellite coatings were advantageous for use in erosion–corrosion environments and can even function at relatively high temperatures. They reported that this is because wear resistance is promoted by the harder complex carbides of chromium and tungsten, while corrosion resistance is enhanced by the presence of cobalt in the matrix. For pure metals, some correlation between erosion rate and hardness has been shown [11]. However, several other observations have shown that the erosion rate is not dependent on material hardness [12,13]. Miller and Coyle [14] studied the erosion resistance of two different powder metallurgy Stellite alloys (3 and 6) where they showed that erosion resistance cannot always be predicted by macro hardness of a material. The effect of microstructure and especially the size of carbides and the abrasion resistance was the focus of a study by Shetty et al. [15]. They examined the effect of carbide size on abrasive wear resistance using a scratch test with Al2 O3 particles of diameter in the range from 300 to 500 m and
a Vickers diamond pyramid. For a structure containing fine carbides (sub 10 m), they observed the complete removal of small carbides with chips of debris. They also demonstrated the influence of work-hardening on the wear resistance. Their results showed that the width of the scratch was reduced by a factor of two, due to the work-hardening of the surface during the pre-abrasion test. In this paper the erosion–corrosion performance of Stellite alloys 6 and 706 in the cast and Hot Isostatically Pressed (HIPed) formed is compared. The paper specifically focuses on the effect of tungsten and molybdenum as alloying elements at approximately 5 wt.% in alloy 6 and 706, respectively. The effect of the alloying elements and processing conditions on the material loss rates and mechanisms in erosion–corrosion is discussed. Because in several applications stainless steels are used as alternative materials to Stellite alloys some comparison is made with an austenitic stainless steel (UNS S31603) and a higher grade super duplex stainless (UNS S32760). The extent of the interactions between corrosion and erosion are determined for a range of conditions and these are discussed in relation to the overall material performance.
2. Experimental details The focus of this work is to assess the tribo-corrosion (specifically erosion–corrosion) behaviour of a cobalt-based alloy Stellite 6 and Stellite 706 manufactured by two different techniques; Hot Isostatic Pressing (HIP) and Investment Casting. For comparative purposes two stainless steels are included: UNS S31603, an austenitic stainless steel having 18% chromium, 10% nickel, and a super duplex stainless steel (UNS S32760) with 25% chromium, 6% molybdenum, 9% nickel. Erosion–corrosion tests were performed using a recirculating impinging jet apparatus as shown in Fig. 1. The dual nozzle system delivered an impinging liquid–solid jet at a liquid velocity of 17 m/s at an angle of 90◦ on to the surface of the sample. The velocities for all tests was 17 m/s chosen to represent the upper range of impeller velocities for pumping applications in the oil and gas industries. The sand used in the tests was a silica sand (Congleton HST 60) and the particles were rounded. The sand size distribution is shown in Table 1. In this study solid loadings in the range of 0–500 mg/l were studied. In each test the solid content was measured (by taking water samples, filtering and weighing the solid residue) at least three times during the test to ensure that the solid loading ejected from the nozzle was within ±50 mg/l of the quoted value. The tests ran for 8 h and after each test the samples were weighed and the extent of the mass loss was determined. This was done using a Mettler Toledo balance with an accuracy of ±0.1 mg. For all conditions tests were repeated at least three times. The mean is plotted and the maximum and minimum values are given in two error bars.
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Fig. 1. Re-circulating impingement rig used for erosion–corrosion tests in liquid–solid slurries.
Hardness measurements were performed using a microhardness Vickers indenter Mitutoyo MVK-H1 with a 100 g applied load. In situ electrochemical tests were facilitated through connection of the working electrode (the sample under erosion–corrosion) to a three-electrode cell. The reference electrode used was Ag/AgCl and the auxiliary electrode used was Pt. Electrochemical tests were performed as described in the following paragraphs. For electrochemical corrosion tests (in static conditions) a wire was soldered on the rear of the sample and it was then embedded in a non-conducting resin. Samples, of 4.5 cm2 surface area, were prepared for corrosion and erosion–corrosion testing by grinding with 600 and 1200 grit SiC paper followed by diamond polishing with 6-micron diamond grit. The surface was rinsed with methanol after polishing and dried with compressed air. For electrochemical connection for erosion–corrosion tests a wire was screwed into the rear of the sample and then the connection was sealed using a shrink fit plastic. The whole configuration was then inserted into a Perspex holder, which isolated the rest of the sample except the face exposed to the impinging jet. Direct current (dc) accelerated test techniques are used to study the susceptibility of the alloys to localised passivity breakdown and the onset of corrosion. In this study dc anodic and cathodic polarisation tests were carried out to assess the kinetics of the anodic and cathodic reactions occurring at the sample surface under specific environment conditions (in
static solution or under the impinging jet). In anodic polarisation tests the potential was shifted from the free corrosion potential (Ecorr ) in the noble (positive potential) direction at a rate of 15 mV/min. Once a current density of 500 A/cm2 was achieved in the external circuit between the sample and the auxiliary electrode the potential scan was reversed and the potential was scanned in the negative direction to reach Ecorr . Light microscopy and scanning electron microscopy (SEM) with energy dispersive X-ray analysis (EDX) attachment were used in this study. The SEM is equipped with a LaB6 gun, and is capable of operating as a conventional highvacuum SEM, or under low vacuum in ESEM mode. The SEM is fully equipped with a range of secondary electron (SE) and back-scattered electron (BSE) detectors. Corrections for atomic number, absorption and fluorescence (ZAF) are achieved through a virtual standard calibration routine. In the initial stages of the programme this was for characterisation of the materials in the as-received conditions. After corrosion and erosion–corrosion tests the samples were examined by SEM to assess the extent of surface degradation by mechanical and electrochemical mechanisms. X-ray diffraction (XRD) was used to analyse the crystalline components of the alloys. A Siemens D500 diffractometer with copper radiation K␣1 + K␣2 and a scintillation counter (point detector), which produces a θ versus 2θ scan (Bragg Brentano geometry) has been used for the analyses.
Table 1 Sand size distribution from sieve analysis of the silica sand (Congleton HST 60) used in the erosion–corrosion tests
3. Results
Sand size (m)
% of total mass
3.1. Material characterisation
<160 180 250 300 >425
16 37 21 20 7
The chemical composition of both samples, measured using flame spectroscopy is given in Table 2. Fig. 2a–d shows the microstructure of the as-polished cast Stellite 6, HIPed Stellite 6, cast 706 and HIPed 706 respectively. Also in
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Table 2 Measured chemical composition of HIPed and cast Stellite 6 and Stellite 706 Alloy
Production method
Elements mass% Co
Cr
W
C
Fe
Ni
Mo
Stellite 6
HIPed Cast
Bal. Bal.
29 27.6
4.4 4.8
1.1 1.1
1.8 0.8
2.5 0.8
0.3 0.6
Stellite 706
HIPed Cast
Bal. Bal.
28.2 28.6
0.1 0.1
1.2 1.3
1.7 1.5
2.4 1.6
5.5 6.1
Tables 3 and 4 the EDX analysis of the different regions on the samples are shown. It can be clearly seen from the back-scattered SEM images that the HIPed material presents a fine and homogeneous carbide distribution compared to the cast alloys due to the nature of the isostatic pressure applied promoting uniform densification [16]. The structure of both HIPed Stellite 6 and Stellite 706 are similar except mainly for the variation in the propor-
tion and the types of carbides. Importantly, Stellite 706 has molybdenum-rich carbides in addition to chromium-rich carbides. The average diameter of the chromium-rich carbides in Stellite 706 is similar to those in Stellite 6 (around 2.2 m) and the average diameter of the molybdenum-rich carbides is around 1.6 m. The measured micro hardness values are summarised in Table 5. Also the volume fractions of carbides in the HIPed alloys Stellite 6 and Stellite 706 are shown in
Fig. 2. Back scattered SEM images of (a) cast Stellite 6, (b) HIPed Stellite 6, (c) cast Stellite 706, (d) HIPed Stellite 706.
U. Malayoglu, A. Neville / Wear 259 (2005) 219–229 Table 3 EDX analysis of the cast and HIPed Stellite 6 Elements wt.%
Co Cr W C Others
Cast Stellite 6
HIPed Stellite 6
A1
A2
B1
B2
13.2 72.7 5.1 8.7 0.3
65.3 23.5 5.1 1.3 4.8
16.4 71.1 3.4 7.6 1.5
60.2 26.8 4.9 0.9 7.2
Table 4 EDX analysis of the cast and HIPed Stellite 706 Elements wt.%
Cast Stellite 706
HIPed Stellite 706
C1
C2
C3
D1
D2
D3
Co Cr Mo C Others
10.6 68.9 8.1 11.6 0.6
46.6 19.7 26.4 3.8 3.2
64.8 24.9 4.1 2.9 3
34.8 46.6 4.3 12.2 1.9
30.7 24.1 29.7 8.3 7
66.8 22.9 3.3 2.6 4.1
Table 5 Micro hardness values of the Stellite alloys Material
Average hardness HV100
Standard deviation
Stellite 6 Cast Stellite 6 HIPed Stellite 706 Cast Stellite 706 HIPed
482 434 456 480
65 24 59 24
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Both cast Stellite 6 and Stellite 706, have dendrites with lamellar eutectic carbides. Due to the differences in the cooling rate the chemical composition of the first part of the dendrite cooling is different than the last part to solidify. This difference in the chemical composition of the dendrite can be clearly seen by examining the EDX results from Fig. 2c in cast Stellite 706. As reported previously for Stellite 6 corrosion initiated and propagated at the interface between the matrix and the carbide where there is a difference in the chemical composition [17]. Under erosion–corrosion conditions if the interface is weakened as a result of corrosion attack this may cause the carbide to pull out easily and accelerate the overall wear rate. The XRD spectra for the as-polished surfaces are shown in Fig. 3 for cast and Fig. 4 for HIPed Stellite 706. When compared with the XRD patterns for cast and HIPed Stellite 6 reported previously [17], it becomes apparent that the samples primarily include a large amount of carbides such as M7 C3 and M23 C6 , which are basically chromium carbides containing cobalt in substitution from chromium and MC carbides, which are defined as refractory element carbides [18] consisting of tungsten and molybdenum. Also both cast and HIPed Stellite 706 have phases like Co3 Mo, Co7 Mo6 which are reported to form due to the excess amount of molybdenum present [19]. 3.2. Erosion–corrosion material loss assessment
Table 6 Carbide volume fractions of the HIPed 6 and 706 Material
Cr rich carbides (%)
Mo rich carbides (%)
Stellite 6 Stellite 706
12.2 15.3
– 5.5
Table 6 and it is clear that the total volume fraction of carbides is greater for Stellite 706. The volume fraction of carbides in the cast alloy was not determined due to inaccuracies which occur when the hard phase has a complex form and where sub-micron size carbides exist.
In Fig. 5a and b, the total weight loss (TWL) measured after exposure to the impinging jet at 17 m/s is shown after 8 h test period for the cast and HIPed alloys at 20 and 50 ◦ C and at two different solid loadings. For comparison, stainless steels UNS S31603 and UNS S32750 are also included. From these results the following observations can be made: • Both Stellite 6 and Stellite 706 in cast and HIPed form show a consistently lower TWL than UNS S31603 and UNS S32750 stainless steels.
Fig. 3. XRD pattern of cast Stellite 706 and the possible phases.
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Fig. 4. XRD pattern of HIPed Stellite 706 and the possible phases.
• The distinction between the TWL of cast and HIPed Stellite 706 is lower than between the cast and HIPed Stellite 6. • TWL of the alloys is affected by temperature and solid loading. Increase of both increased the overall erosion–corrosion rate. • Among the Stellite alloys, HIPed Stellite 706 had the lowest TWL in all test conditions. Fig. 6 shows the different regions forming on the surface during erosion–corrosion conditions. Beneath the water jet, the surface of specimen can be divided into three
zones due to the fluid dynamic influence. The most frequent and highest angle solid impacts occur right beneath the water jet as shown in zone I and the diameter of this zone, central worn region, is equal to the diameter of impinging jet (4 mm). Outside the central region (zone II) there is a lower frequency impact zone, named as “halo” region by Wood [20]. At the edge of the sample (zone III), few solids impact the surface and the loss of thickness is not measurable. Preferential removal of the cobalt-rich matrix in the low angle impact zone was shown for cast Stellite 6 by using the atomic force microscope [5]. The matrix was extruded out by the low angle impacts and the carbides were left standing out from the surface. The same kind of preferential removal of the matrix in cast Stellite 706 is also seen as shown in Figs. 7 and 8. On the HIPed sample, there is no evidence of discrete carbides and the whole surface shows general plastic deformation (Fig. 9). 3.3. Corrosion measurements—in situ
Fig. 5. Total weight loss measurements: (a) 18 ◦ C and (b) 50 ◦ C, respectively.
Figs. 10 and 11 show the anodic polarisation graphs for cast and HIPed Stellite 6 under liquid–solid erosion– corrosion conditions at 20 ◦ C and 50 ◦ C for 200 mg/l solid loading condition. As reported earlier HIPed Stellite 6 possesses superior corrosion resistance at both temperatures compared to cast Stellite 6 [17]. In the case of Stellite 706 as shown in Figs. 12 and 13 both cast and HIPed materials had a higher breakdown potential in static conditions than Stellite 6 showing a better resistance to passivity breakdown. Passivity is a key issue in erosion–corrosion of passive alloys since the rate at which the passive film reforms, once it has been ruptured, has a strong influence on the corrosion–erosion interactions [22]. Fig. 12 also shows the comparison of the anodic polarisation curves under static conditions and under liquid–solid impingement conditions. Although the form of anodic polarisation curve appears similar, the current in the stable region
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Fig. 6. White light interferometry images of HIPed Stellite 706 showing the different regions on the sample under erosion–corrosion conditions.
between potential values of −100 mV to 800 mV is far in excess of what would be expected to indicate passivity. There are three distinct regions defined as region I (active region) region II (pseudo-passive region) and region III (breakdown region). On increasing the potential from Ecorr in the positive direction the current density i increases. The increase in current density, which is as shown in Fig. 12 region I indicates that rapid dissolution occurs in the wear zone and again a re-passivation in the wear effected zone. The rate of this reaction taking place in the areas where the passive film is removed, initially increases with potential. Similar characteristics have been observed for stainless steels [22] and cur-
rent in region II is determined by the level of solid loading, the temperature and material characteristics. It represents the steady state charge transfer in the zone activated by the impacting solids. In region III of the anodic polarisation curve the passive areas of the sample break down through initiation of localised attack. This potential (Eb ) is independent of solid loading (Figs. 12 and 13) but is affected by temperature (Figs. 10 and 11). The corrosion current densities (icorr ) of the four materials at two different sand loading and two different test temperatures are presented in Table 7. Judging from the corrosion current density values the corrosion resistance of these specimens under erosion–corrosion conditions can be ranked in
Fig. 7. White light interferometry images of cast Stellite 706 showing preferential dissolution of matrix in zone II. The carbides protruding from the surface cause the rough appearance of the perimeter of the wear scar.
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Fig. 10. Anodic polarisation curves of cast Stellite 6 at 200 mg/l solid loading at 20 ◦ C and 50 ◦ C at 17 m/s in 3.5% NaCl solution.
Fig. 8. SEM image of the low angle impact zone at 50 ◦ C 500 mg/l solid loading conditions showing the carbide network protruding on the surface of cast Stellite 706.
Fig. 11. Anodic polarisation curves of HIPed Stellite 6, at 200 mg/l solid loading at 20 ◦ C and 50 ◦ C at 17 m/s in 3.5% NaCl solution.
descending order as follows: HIPed Stellite 706 > Cast Stellite 706 > HIPed Stellite 6 > Cast Stellite 6.
4. Discussion
Fig. 9. Secondary electron SEM image active wear zone (zone I) on the surface of HIPed Stellite 706 showing heavy plastic deformation at 50 ◦ C, 500 mg/l solid loading conditions.
The loss of material by erosion–corrosion is a complex process affected by many interacting variables. Thus changing any variable can have an important consequence on the behavioural pattern of the total system. In this paper two different cobalt-based super alloys, Stellite 6 and Stellite 706
Fig. 12. Effect of solid loading on the anodic polarisation of HIPed Stellite 706 at 20 ◦ C compared with the anodic polarisation curve under static conditions at 17 m/s in 3.5% NaCl solution.
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Table 7 Corrosion current densities (icorr ) values (A/cm2 ) at two different sand loading ant two different temperature Cast Stellite 6 20 ◦ C 50 ◦ C
HIPed Stellite 6
Cast Stellite 706
HIPed Stellite 706
200 mg/l
500 mg/l
200 mg/l
500 mg/l
200 mg/l
500 mg/l
200 mg/l
500 mg/l
7 12
13 33
4 8
10 18
3 8
10 19
3 6
8 16
produced by two different methods (casting and HIPing) have been subjected to liquid–solid erosion–corrosion where the impact velocity, erodent characteristics and impact angle remain constant. Erodent concentration and temperature of the slurry were changed. The effect of increase in temperature and increase in solid loading were discussed in detail in the previous work [5]. The difference in the erosion–corrosion resistance of the two cast and HIPed alloys are primarily due to the differences in: (a) microstructure – which affects mechanical integrity; (b) corrosion behaviour – which is important for the formation of the passive film and the extent of charge transfer during depassivation; (c) strength of matrix (affected by alloying elements) – which affects mechanical integrity and resistance to charge transfer. In terms of the microstructure effects, the cast samples of both materials experienced higher weight loss due to mechanical erosion compared to HIPed samples because they had larger matrix areas between the dendrites that were unprotected by the harder carbides. The coarser carbide size meant that, when carbide removal did not occur, large pieces of matrix were removed, which also accelerated the wear rate. Under erosion–corrosion conditions these unprotected matrix areas also plays an important role as more matrix area will be in contact with the corrosive environment. To understand why the alloy containing molybdenum has better erosion–corrosion resistance compared to the alloy that contains tungsten the three factors listed above as (a), (b) and (c) are further considered.
Fig. 13. Effect of solid loading on the anodic polarisation of cast Stellite 706 at 20 ◦ C compared with anodic polarisation in static conditions.
When the SEM images of Stellite 706 (both cast and HIPed) and Stellite 6 (both cast and HIPed) (Fig. 2) are compared it can be clearly seen that in addition to the formation of chromium-rich carbides in the alloys with molybdenum there is a formation of a molybdenum-rich carbide. The formation of these secondary carbides depends on the solubility limits of molybdenum and tungsten in cobalt. The solubility limit of both molybdenum and tungsten in cobalt extends to a maximum of 40% by wt. at the eutectic temperature [19]. As molybdenum (atomic weight = 94.94 g) has twice as many atoms as tungsten (atomic weight = 183.84 g) for the same weight percentage, there will be an excessive atomic concentration of molybdenum when a comparable wt.% is added. Because of this the propensity for molybdenum-rich carbides to form is higher. The main role of both tungsten and molybdenum in cobalt-based alloys is in solid-solution strengthening. In Stellite 6 as all the tungsten added takes part in solid solution strengthening, there is no excessive tungsten to form secondary tungsten-rich carbides and, as such, no tungsten-rich carbide formation is observed. As summarised in Table 6 carbide volume fractions on HIPed Stellite 706 are much higher compared to the HIPed Stellite 6. The volume fraction of carbide directly affects the erosion rate of the alloy and is one major contributory factor to Stellite 706 having superior erosion–corrosion resistance. As summarised in Table 3, the microhardness of cast Stellite 6 is higher compared to HIPed Stellite 6, but the erosion–corrosion resistance of the HIPed alloy is higher (Fig. 5). Again comparison of the both cast alloys of Stellite 6 and Stellite 706 (Fig. 5), the alloy with lower hardness value has higher erosion–corrosion resistance which shows that there is no direct relation between the hardness and erosion–corrosion resistance which is not unexpected if the controlling factors in erosion–corrosion are considered. In erosion alone little correlation between hardness and wear rate is seen partly due to the ideas of Hutchings and Levy [21] who postulated that the temperature rise during particle impact “anneals out” the effects of the hardening mechanism. Surface hardness after erosion might be a useful parameter to rank the erosion resistance of the materials. In addition in erosion–corrosion there is a substantial effect of corrosion which is unlikely to be directly linked to hardness and is more likely to be affected by depassivation/repassivation characteristics. Since the shape, size and the composition of the carbides are different in both cast and HIPed microstructure, support from the matrix is essential. In the case of abrasion wear it has been shown that sufficient support by the matrix can prevent
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Table 8 Normalised alloy content of Stellite 6 and Stellite 706 Alloy Cast Stellite6 HIPed Stellite 6 Cast Stellite 706 HIPed Stellite 706
Normalised alloy content (NAC) 1 1.2 1.36 1.39
pulling out of the carbides [23]. This is also the case where sand particles impact at low angles, as is the case in zone II in Fig. 6. The strength of the matrix is considered in detail here. Kosel et al. [24] showed the effect of the matrix strength can affect the wear rate. They used the normalised alloy content (NAC) as a measure of the matrix strength. They defined the normalised alloy content as the sum of the weight percentages of nickel, vanadium, tungsten; the alloys that are used as solid-solution strengtheners to give an approximate measure of the degree of solid-solution strengthening. By using the same approach the NAC of the alloys Stellite 706 and 6 were calculated and normalised against cast Stellite 6 as shown in Table 8. As shown in Fig. 14 there is linear relationship between the NAC and the erosion–corrosion resistance of the alloys. The alloys with high NAC gave a lower weight loss in all test conditions. This strain hardening of the matrix by alloy addition is explained through the change in the stacking fault energy (SFE). Excellent erosion resistance of cobalt-based Stellite alloy is greatly attributed to the low SFE. Molybdenum, tungsten and chromium tend to stabilize the hcp allotrope and decrease the SFE [18]. The lower the SFE (the greater the width of stacking fault) the more difficult cross-slip is and the higher the work-hardening and the strain to fracture. The materials with low SFE tend to strain harden rapidly, and show the highest erosion resistance. Both molybdenum and tungsten lowers the SFE, and so it is likely that in both Stellite 6 and 706 there will be some effect of SFE reduction. However the exact degree of this effect still needs further investigation. In terms of the corrosion behaviour under liquid–solid impingement, generation of fresh metal surface and the repassivation ability are two important factors particularly in the
Fig. 14. Total weight loss test results plotted vs. normalised alloy content.
case of discontinuous impacts of solid particles where depassivation caused by solid particle impacts is followed by a process of new surface repassivation. The features of anodic polarisation data can be very useful in determining the nature of the liquid–solid particle interactions with the solid surface and the associated corrosion activity. Both cast and HIPed Stellite 706 showed a pseudo passive region (region II in Fig. 12) in which the measured current did not change with the potential but where there is clear charge transfer and loss of true passivity. This indicates that the surface protective film is continuously destroyed by the mechanical factors but that the degree of charge transfer associated with the impacts is lower than on Stellite 6. Comparison of the corrosion current densities in Table 7 shows that in all conditions HIPed Stellite 706 is lowest which in turn means the resistance to breakdown of the passive film is highest. This can be due to the fact that it contains more carbides so that less matrix area is in contact with the corrosive liquid and the chemical composition of the matrix contains more corrosion-resistant elements like chromium to maintain integrity against charge transfer. Also comparison of the is (which is the current value at the start of the pseudo-passive region) as shown in Fig. 15, shows that the same trend in icorr is observed. A higher icorr (corrosion in the active state) also translates to a higher is (charge transfer associated with depassivation regions) such that the resistance of HIPed Stellite 706 is greatest and cast Stellite 6 lowest. Comparison of the two conditions (Fig. 15a
Fig. 15. Comparison of the anodic polarisation curves of Stellite 6 and Stellite 706 at (a) 18 ◦ C, 200 mg/l and (b) 50 ◦ C, 500 mg/l.
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and b) show that increasing the solid loading and temperature, increases is since more charge transfer occurs due to the impacts.
5. Conclusions The erosion–corrosion of cast and HIPed Stellite 706 has been examined in aggressive erosion–corrosion conditions and compared with cast and HIPed Stellite 6 alloy. The following conclusions are reached. All Stellite alloys for the conditions studied have superior resistance to superduplex (UNS S32760) and austenitic (UNS S31603) stainless steels. HIPed alloys show improved resistance to cast alloys but the degree of improvement is less in Stellite 706 than Stellite 6. 1. Microstructure plays a crucial role in the erosion– corrosion performance of the alloys. Addition of Mo is much more effective than W addition in conferring erosion–corrosion resistance to cobalt-based alloys due to formation of secondary carbides. 2. The strength of matrix, which is affected by alloying elements, is also to be considered as an important factor to affect the erosion–corrosion performance of the alloys. 3. The rate of the charge transfer during depassivation and the rate of repassivation of the surface is affected by the microstructure and alloying elements in the matrix. Acknowledgements The authors acknowledge the financial support to UM from Deloro Stellite and Heriot-Watt University. References [1] X. Hu, A. Neville, An examination of the electrochemical characteristics of two stainless steels (UNS S32654 and UNS S31603) under liquid–solid impingement, Wear 256 (2004) 537–544. [2] D.L. Klastrom, Wrought cobalt-based superalloys, J. Mater. Eng. Perform. 2 (1993) 523–530. [3] A. Neville, M. Reyes, T. Hodgkiess, A. Gledhill, Mechanisms of wear on a Co-based alloy in liquid–solid slurries, Wear 238 (2000) 138–150. [4] A. Neville, T. Hodgkiess, Characterisation of high-grade alloy behaviour in severe erosion–corrosion conditions, Wear 233–235 (1999) 596–607.
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[5] U. Malayoglu, A. Neville, Comparing the performance of HIPed and cast Stellite 6 alloy in liquid–solid slurries, Wear 255 (2003) 181–194. [6] J.C. Shin, J.M. Doh, J.K. Yoon, L. Dok-Yol, J.S. Kim, Effect of molybdenum on the microstructure and wear resistance of cobaltbased Stellite hardfacing alloys, Surf. Coating Technol. 166 (2003) 117–126. [7] P. Crook, The effect of dilution upon the corrosion and wear properties of cobalt-based weld overlays, Corr. Sci. 35 (1993) 647– 653. [8] C.J. Heathcock, A. Ball, Cavitations erosion of cobalt-based Stellite alloys, cemented carbides and surface-treated low alloy steels, Wear 74 (1982) 11–26. [9] M.K. Lee, W. Kim, K.C. Rhee, J.W. Lee, Investigation of liquid impact erosion for 12 Cr steel and Stellite 6 B, J. Nuclear Mater. 257 (1998) 134–144. [10] M. Wong-Kian, L.A. Cornish, A. Van Bennekom, Comparison of erosion–corrosion behaviour of Hot Iso-Statically pressed and welded Stellite coatings, J. South Afr. Inst. Mining Metall. Nov/Dec (1995) 319–335. [11] I. Finnie, J. Wolak, Y. Kabil, Erosion of metals by solid particles, J. Mater. 2 (1967) 682–700. [12] A. Levy, S. Jahanmir, The effect of microstructure of ductile alloys on solid particle erosion behavior, in: K. Natesan (Ed.), Corrosion–Erosion Behavior of Materials, The Metallurgical Society of AIME, Warrendale, PA, 1980, pp. 177–189. [13] B.F. Levin, J.N. Dupont, A.R. Marder, Weld overlay coatings for erosion control, Wear 181–183 (1995) 810–820. [14] A.E. Miller, J.P. Coyle, An orifice erosion test sensitive to metallurgical microstructure, Wear 47 (1978) 211–214. [15] H.R. Shetty, T.H. Kosel, N.F. Fiore, A study of abrasive wear mechanisms using diamond and alumina scratch tests, Wear 80 (1982) 347–376. [16] D.J. Stephenson, M. Downing, Hot isostatic pressing of PM and cast components, Mater. High Temp. 9 (1991) 59–68. [17] U. Malayoglu, A. Neville, Assessing the kinetics and mechanisms of corrosion of cast and HIPed Stellite 6 in aqueous saline environments, Corr. Sci., in press. [18] The Superalloys, J. Wiley & Sons, New York, London, Sydney, Toronto, 1984. [19] F.M. Ghaini, Ph.D. thesis, Cranfield Institute of Technology, 1990. [20] F.W. Wood, Erosion by solid-particle impacts—a testing update, J. Test. Eval. 14 (1986) 23–27. [21] I. Hutchings, A.V. Levy, Thermal effects in the erosion of ductile metals, Wear 131 (1989) 105–121. [22] A. Neville, X. Hu, Mechanical and electrochemical interactions during liquid–solid impingement on high-alloy stainless steels, Wear 251 (2001) 1284–1294. [23] V.M. Desai, C.M. Rao, T.H. Kosel, F. Fioren, Effect of carbide size on the abrasion of cobalt base powder metallurgy alloys, Wear 94 (1984) 89–101. [24] T.H. Kosel, N.F. Fiore, J.P. Coyle, S.P. Udvargy, W.A. Konkel, Microstructure and wear in high alloys, in: Proceedings of Symposium on Corrosion–Erosion Behavior of Materials, Metallurgical Society of AIME, Warrendale, PA, 1980, pp. 190–208.
Mo and W as alloying elements in Co-based alloys—their effects on erosion–corrosion resistance Ugur Malayoglu, Anne Neville ∗ Corrosion and Surface Engineering Research Group, School of Mechanical Engineering, University of Leeds, Leeds LS2 9JT, UK Received 27 July 2004; received in revised form 27 January 2005; accepted 2 February 2005 Available online 10 May 2005
Abstract In this paper results from erosion–corrosion tests performed under liquid–solid erosion conditions in 3.5% NaCl liquid medium are reported. The focus of the paper is to compare the behaviour of Stellite alloy 706 which contains 5 wt.% molybdenum with Stellite alloy 6 which contains 4.8 wt.% tungsten, both in cast and Hot Isostatically Pressed (HIPed) form, in terms of their electrochemical corrosion characteristics, their resistance to mechanical degradation and relationship between microstructure and degradation mechanisms. It has been shown that both cast and HIPed Stellite 706 possess better erosion and erosion–corrosion resistance than Stellite 6 counterparts under a wide range of conditions. The significant role of the microstructure and specifically the type of carbides in affecting the erosion–corrosion performance of the alloys is demonstrated and discussed. Also it has been shown that in these multiphase alloys there is no direct relationship between the hardness and erosion–corrosion resistance. © 2005 Elsevier B.V. All rights reserved. Keywords: Co-based alloys; Liquid–solid slurries; Erosion; Corrosion
1. Introduction In the chemical, petrochemical and pump industries, machine parts often work in conditions where erosion and corrosion processes, acting together, are the main failure mechanisms. The deleterious synergistic effect of the erosion removes either corrosion product or the passive layer that is protecting the underlying surface. When a passive layer is removed, the time taken for it to repassivate is an important consideration in the assessment of wear rates. The rate of repassivation determines the amount of charge that can transfer when the surface is activated by an erosion event (e.g. impact of sand). The rate of repassivation in relation to the frequency of impact is an important consideration in erosion–corrosion. On stainless steels [1] it was shown that higher amounts of key elements (chromium, molybdenum) can lead a faster repassivation during erosion impacts.
∗
Corresponding author. Tel.: +44 113 343 6812; fax: +44 113 242 4611. E-mail address: [email protected] (A. Neville).
0043-1648/$ – see front matter © 2005 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2005.02.038
Cobalt-based alloys have enjoyed extensive use in wearrelated engineering applications for well over 50 years because of their inherent high-strength, corrosion resistance and ability to retain hardness at elevated temperatures [2]. In recent years a concentrated effort has been made to understand the deformation characteristics of cobalt-based alloys exposed to erosion–corrosion environments in order to optimize those factors contributing to their erosion resistance [3–5]. Alloying cobalt with chromium and various quantities of carbon, tungsten and molybdenum produces a family of alloys which can have excellent resistance to corrosion and/or erosion. Understanding how microstructural changes, as a result of alloying, affect corrosion and erosion resistance is critical to optimising the alloy for a particular purpose. In cobalt-based alloys, the key element chromium is added in the range of 20–30 wt.% to improve corrosion and impart some measure of solid-solution strengthening. Where carbide precipitation strengthening is a desirable feature, chromium also plays a strong role through the formation of a series of varying chromium–carbon ratio carbides such as M7 C3 and M23 C6 . Alloying elements like tungsten, molybdenum and
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tantalum are added to cobalt for solid solution strengthening. If these metals are added in excess of their solubility, formation of carbides like MC and M6 C is likely to occur. Shin et al. [6] investigated the effect of molybdenum on the microstructure and wear resistance properties of Stellite 6 hardfacing alloy. They showed that with an increase in molybdenum content, the M23 C6 and M6 C type carbides were formed instead of chromium-rich M7 C3 . They concluded that this microstructural change was responsible for the improvement of the mechanical properties such as hardness and wear resistance of molybdenum-modified Stellite 6 hardfacing alloy. Most cobalt-based alloys possess outstanding cavitation resistance compared to stainless steels which has been shown to be independent of the carbon content (hence hardness), and has been attributed by Crook [7] to crystallographic transformation, under stress, from the face centred cubic (fcc) to hexagonal close packed (hcp) structure by twinning. Heathcock and Ball [8] studied the cavitation resistance of a number of Stellite alloys, cemented carbides and surface-treated alloy steels and showed that in Stellite alloys the cobalt-rich solid solution, incorporating elements such as chromium, tungsten and molybdenum is highly resistant to erosion due to a rapid increase in the work-hardening rate and the strain to fracture which are caused by deformation twinning. Lee et al. [9] compared the liquid impact erosion resistance of 12 Cr steel with a Vickers hardness of 380 kg/mm2 (∼39 MPa) and Stellite 6B with a hardness value of 420 kg/mm2 (∼43 MPa). The liquid impact erosion resistance of Stellite 6B was at least six times greater than that of 12 Cr steel, implying that hardness is not the governing factor for liquid erosion. Stellite 6B also showed very different behaviour in liquid impact erosion in comparison with 12 Cr steel. They concluded that the superior erosion resistance of Stellite 6B results from the cobalt matrix whose deformation appeared mostly as mechanical twins and the material removal was more dominant in the hard carbide precipitates than in the ductile cobalt matrix. Wong-Kian [10] showed that Stellite coatings were advantageous for use in erosion–corrosion environments and can even function at relatively high temperatures. They reported that this is because wear resistance is promoted by the harder complex carbides of chromium and tungsten, while corrosion resistance is enhanced by the presence of cobalt in the matrix. For pure metals, some correlation between erosion rate and hardness has been shown [11]. However, several other observations have shown that the erosion rate is not dependent on material hardness [12,13]. Miller and Coyle [14] studied the erosion resistance of two different powder metallurgy Stellite alloys (3 and 6) where they showed that erosion resistance cannot always be predicted by macro hardness of a material. The effect of microstructure and especially the size of carbides and the abrasion resistance was the focus of a study by Shetty et al. [15]. They examined the effect of carbide size on abrasive wear resistance using a scratch test with Al2 O3 particles of diameter in the range from 300 to 500 m and
a Vickers diamond pyramid. For a structure containing fine carbides (sub 10 m), they observed the complete removal of small carbides with chips of debris. They also demonstrated the influence of work-hardening on the wear resistance. Their results showed that the width of the scratch was reduced by a factor of two, due to the work-hardening of the surface during the pre-abrasion test. In this paper the erosion–corrosion performance of Stellite alloys 6 and 706 in the cast and Hot Isostatically Pressed (HIPed) formed is compared. The paper specifically focuses on the effect of tungsten and molybdenum as alloying elements at approximately 5 wt.% in alloy 6 and 706, respectively. The effect of the alloying elements and processing conditions on the material loss rates and mechanisms in erosion–corrosion is discussed. Because in several applications stainless steels are used as alternative materials to Stellite alloys some comparison is made with an austenitic stainless steel (UNS S31603) and a higher grade super duplex stainless (UNS S32760). The extent of the interactions between corrosion and erosion are determined for a range of conditions and these are discussed in relation to the overall material performance.
2. Experimental details The focus of this work is to assess the tribo-corrosion (specifically erosion–corrosion) behaviour of a cobalt-based alloy Stellite 6 and Stellite 706 manufactured by two different techniques; Hot Isostatic Pressing (HIP) and Investment Casting. For comparative purposes two stainless steels are included: UNS S31603, an austenitic stainless steel having 18% chromium, 10% nickel, and a super duplex stainless steel (UNS S32760) with 25% chromium, 6% molybdenum, 9% nickel. Erosion–corrosion tests were performed using a recirculating impinging jet apparatus as shown in Fig. 1. The dual nozzle system delivered an impinging liquid–solid jet at a liquid velocity of 17 m/s at an angle of 90◦ on to the surface of the sample. The velocities for all tests was 17 m/s chosen to represent the upper range of impeller velocities for pumping applications in the oil and gas industries. The sand used in the tests was a silica sand (Congleton HST 60) and the particles were rounded. The sand size distribution is shown in Table 1. In this study solid loadings in the range of 0–500 mg/l were studied. In each test the solid content was measured (by taking water samples, filtering and weighing the solid residue) at least three times during the test to ensure that the solid loading ejected from the nozzle was within ±50 mg/l of the quoted value. The tests ran for 8 h and after each test the samples were weighed and the extent of the mass loss was determined. This was done using a Mettler Toledo balance with an accuracy of ±0.1 mg. For all conditions tests were repeated at least three times. The mean is plotted and the maximum and minimum values are given in two error bars.
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Fig. 1. Re-circulating impingement rig used for erosion–corrosion tests in liquid–solid slurries.
Hardness measurements were performed using a microhardness Vickers indenter Mitutoyo MVK-H1 with a 100 g applied load. In situ electrochemical tests were facilitated through connection of the working electrode (the sample under erosion–corrosion) to a three-electrode cell. The reference electrode used was Ag/AgCl and the auxiliary electrode used was Pt. Electrochemical tests were performed as described in the following paragraphs. For electrochemical corrosion tests (in static conditions) a wire was soldered on the rear of the sample and it was then embedded in a non-conducting resin. Samples, of 4.5 cm2 surface area, were prepared for corrosion and erosion–corrosion testing by grinding with 600 and 1200 grit SiC paper followed by diamond polishing with 6-micron diamond grit. The surface was rinsed with methanol after polishing and dried with compressed air. For electrochemical connection for erosion–corrosion tests a wire was screwed into the rear of the sample and then the connection was sealed using a shrink fit plastic. The whole configuration was then inserted into a Perspex holder, which isolated the rest of the sample except the face exposed to the impinging jet. Direct current (dc) accelerated test techniques are used to study the susceptibility of the alloys to localised passivity breakdown and the onset of corrosion. In this study dc anodic and cathodic polarisation tests were carried out to assess the kinetics of the anodic and cathodic reactions occurring at the sample surface under specific environment conditions (in
static solution or under the impinging jet). In anodic polarisation tests the potential was shifted from the free corrosion potential (Ecorr ) in the noble (positive potential) direction at a rate of 15 mV/min. Once a current density of 500 A/cm2 was achieved in the external circuit between the sample and the auxiliary electrode the potential scan was reversed and the potential was scanned in the negative direction to reach Ecorr . Light microscopy and scanning electron microscopy (SEM) with energy dispersive X-ray analysis (EDX) attachment were used in this study. The SEM is equipped with a LaB6 gun, and is capable of operating as a conventional highvacuum SEM, or under low vacuum in ESEM mode. The SEM is fully equipped with a range of secondary electron (SE) and back-scattered electron (BSE) detectors. Corrections for atomic number, absorption and fluorescence (ZAF) are achieved through a virtual standard calibration routine. In the initial stages of the programme this was for characterisation of the materials in the as-received conditions. After corrosion and erosion–corrosion tests the samples were examined by SEM to assess the extent of surface degradation by mechanical and electrochemical mechanisms. X-ray diffraction (XRD) was used to analyse the crystalline components of the alloys. A Siemens D500 diffractometer with copper radiation K␣1 + K␣2 and a scintillation counter (point detector), which produces a θ versus 2θ scan (Bragg Brentano geometry) has been used for the analyses.
Table 1 Sand size distribution from sieve analysis of the silica sand (Congleton HST 60) used in the erosion–corrosion tests
3. Results
Sand size (m)
% of total mass
3.1. Material characterisation
<160 180 250 300 >425
16 37 21 20 7
The chemical composition of both samples, measured using flame spectroscopy is given in Table 2. Fig. 2a–d shows the microstructure of the as-polished cast Stellite 6, HIPed Stellite 6, cast 706 and HIPed 706 respectively. Also in
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Table 2 Measured chemical composition of HIPed and cast Stellite 6 and Stellite 706 Alloy
Production method
Elements mass% Co
Cr
W
C
Fe
Ni
Mo
Stellite 6
HIPed Cast
Bal. Bal.
29 27.6
4.4 4.8
1.1 1.1
1.8 0.8
2.5 0.8
0.3 0.6
Stellite 706
HIPed Cast
Bal. Bal.
28.2 28.6
0.1 0.1
1.2 1.3
1.7 1.5
2.4 1.6
5.5 6.1
Tables 3 and 4 the EDX analysis of the different regions on the samples are shown. It can be clearly seen from the back-scattered SEM images that the HIPed material presents a fine and homogeneous carbide distribution compared to the cast alloys due to the nature of the isostatic pressure applied promoting uniform densification [16]. The structure of both HIPed Stellite 6 and Stellite 706 are similar except mainly for the variation in the propor-
tion and the types of carbides. Importantly, Stellite 706 has molybdenum-rich carbides in addition to chromium-rich carbides. The average diameter of the chromium-rich carbides in Stellite 706 is similar to those in Stellite 6 (around 2.2 m) and the average diameter of the molybdenum-rich carbides is around 1.6 m. The measured micro hardness values are summarised in Table 5. Also the volume fractions of carbides in the HIPed alloys Stellite 6 and Stellite 706 are shown in
Fig. 2. Back scattered SEM images of (a) cast Stellite 6, (b) HIPed Stellite 6, (c) cast Stellite 706, (d) HIPed Stellite 706.
U. Malayoglu, A. Neville / Wear 259 (2005) 219–229 Table 3 EDX analysis of the cast and HIPed Stellite 6 Elements wt.%
Co Cr W C Others
Cast Stellite 6
HIPed Stellite 6
A1
A2
B1
B2
13.2 72.7 5.1 8.7 0.3
65.3 23.5 5.1 1.3 4.8
16.4 71.1 3.4 7.6 1.5
60.2 26.8 4.9 0.9 7.2
Table 4 EDX analysis of the cast and HIPed Stellite 706 Elements wt.%
Cast Stellite 706
HIPed Stellite 706
C1
C2
C3
D1
D2
D3
Co Cr Mo C Others
10.6 68.9 8.1 11.6 0.6
46.6 19.7 26.4 3.8 3.2
64.8 24.9 4.1 2.9 3
34.8 46.6 4.3 12.2 1.9
30.7 24.1 29.7 8.3 7
66.8 22.9 3.3 2.6 4.1
Table 5 Micro hardness values of the Stellite alloys Material
Average hardness HV100
Standard deviation
Stellite 6 Cast Stellite 6 HIPed Stellite 706 Cast Stellite 706 HIPed
482 434 456 480
65 24 59 24
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Both cast Stellite 6 and Stellite 706, have dendrites with lamellar eutectic carbides. Due to the differences in the cooling rate the chemical composition of the first part of the dendrite cooling is different than the last part to solidify. This difference in the chemical composition of the dendrite can be clearly seen by examining the EDX results from Fig. 2c in cast Stellite 706. As reported previously for Stellite 6 corrosion initiated and propagated at the interface between the matrix and the carbide where there is a difference in the chemical composition [17]. Under erosion–corrosion conditions if the interface is weakened as a result of corrosion attack this may cause the carbide to pull out easily and accelerate the overall wear rate. The XRD spectra for the as-polished surfaces are shown in Fig. 3 for cast and Fig. 4 for HIPed Stellite 706. When compared with the XRD patterns for cast and HIPed Stellite 6 reported previously [17], it becomes apparent that the samples primarily include a large amount of carbides such as M7 C3 and M23 C6 , which are basically chromium carbides containing cobalt in substitution from chromium and MC carbides, which are defined as refractory element carbides [18] consisting of tungsten and molybdenum. Also both cast and HIPed Stellite 706 have phases like Co3 Mo, Co7 Mo6 which are reported to form due to the excess amount of molybdenum present [19]. 3.2. Erosion–corrosion material loss assessment
Table 6 Carbide volume fractions of the HIPed 6 and 706 Material
Cr rich carbides (%)
Mo rich carbides (%)
Stellite 6 Stellite 706
12.2 15.3
– 5.5
Table 6 and it is clear that the total volume fraction of carbides is greater for Stellite 706. The volume fraction of carbides in the cast alloy was not determined due to inaccuracies which occur when the hard phase has a complex form and where sub-micron size carbides exist.
In Fig. 5a and b, the total weight loss (TWL) measured after exposure to the impinging jet at 17 m/s is shown after 8 h test period for the cast and HIPed alloys at 20 and 50 ◦ C and at two different solid loadings. For comparison, stainless steels UNS S31603 and UNS S32750 are also included. From these results the following observations can be made: • Both Stellite 6 and Stellite 706 in cast and HIPed form show a consistently lower TWL than UNS S31603 and UNS S32750 stainless steels.
Fig. 3. XRD pattern of cast Stellite 706 and the possible phases.
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Fig. 4. XRD pattern of HIPed Stellite 706 and the possible phases.
• The distinction between the TWL of cast and HIPed Stellite 706 is lower than between the cast and HIPed Stellite 6. • TWL of the alloys is affected by temperature and solid loading. Increase of both increased the overall erosion–corrosion rate. • Among the Stellite alloys, HIPed Stellite 706 had the lowest TWL in all test conditions. Fig. 6 shows the different regions forming on the surface during erosion–corrosion conditions. Beneath the water jet, the surface of specimen can be divided into three
zones due to the fluid dynamic influence. The most frequent and highest angle solid impacts occur right beneath the water jet as shown in zone I and the diameter of this zone, central worn region, is equal to the diameter of impinging jet (4 mm). Outside the central region (zone II) there is a lower frequency impact zone, named as “halo” region by Wood [20]. At the edge of the sample (zone III), few solids impact the surface and the loss of thickness is not measurable. Preferential removal of the cobalt-rich matrix in the low angle impact zone was shown for cast Stellite 6 by using the atomic force microscope [5]. The matrix was extruded out by the low angle impacts and the carbides were left standing out from the surface. The same kind of preferential removal of the matrix in cast Stellite 706 is also seen as shown in Figs. 7 and 8. On the HIPed sample, there is no evidence of discrete carbides and the whole surface shows general plastic deformation (Fig. 9). 3.3. Corrosion measurements—in situ
Fig. 5. Total weight loss measurements: (a) 18 ◦ C and (b) 50 ◦ C, respectively.
Figs. 10 and 11 show the anodic polarisation graphs for cast and HIPed Stellite 6 under liquid–solid erosion– corrosion conditions at 20 ◦ C and 50 ◦ C for 200 mg/l solid loading condition. As reported earlier HIPed Stellite 6 possesses superior corrosion resistance at both temperatures compared to cast Stellite 6 [17]. In the case of Stellite 706 as shown in Figs. 12 and 13 both cast and HIPed materials had a higher breakdown potential in static conditions than Stellite 6 showing a better resistance to passivity breakdown. Passivity is a key issue in erosion–corrosion of passive alloys since the rate at which the passive film reforms, once it has been ruptured, has a strong influence on the corrosion–erosion interactions [22]. Fig. 12 also shows the comparison of the anodic polarisation curves under static conditions and under liquid–solid impingement conditions. Although the form of anodic polarisation curve appears similar, the current in the stable region
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Fig. 6. White light interferometry images of HIPed Stellite 706 showing the different regions on the sample under erosion–corrosion conditions.
between potential values of −100 mV to 800 mV is far in excess of what would be expected to indicate passivity. There are three distinct regions defined as region I (active region) region II (pseudo-passive region) and region III (breakdown region). On increasing the potential from Ecorr in the positive direction the current density i increases. The increase in current density, which is as shown in Fig. 12 region I indicates that rapid dissolution occurs in the wear zone and again a re-passivation in the wear effected zone. The rate of this reaction taking place in the areas where the passive film is removed, initially increases with potential. Similar characteristics have been observed for stainless steels [22] and cur-
rent in region II is determined by the level of solid loading, the temperature and material characteristics. It represents the steady state charge transfer in the zone activated by the impacting solids. In region III of the anodic polarisation curve the passive areas of the sample break down through initiation of localised attack. This potential (Eb ) is independent of solid loading (Figs. 12 and 13) but is affected by temperature (Figs. 10 and 11). The corrosion current densities (icorr ) of the four materials at two different sand loading and two different test temperatures are presented in Table 7. Judging from the corrosion current density values the corrosion resistance of these specimens under erosion–corrosion conditions can be ranked in
Fig. 7. White light interferometry images of cast Stellite 706 showing preferential dissolution of matrix in zone II. The carbides protruding from the surface cause the rough appearance of the perimeter of the wear scar.
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Fig. 10. Anodic polarisation curves of cast Stellite 6 at 200 mg/l solid loading at 20 ◦ C and 50 ◦ C at 17 m/s in 3.5% NaCl solution.
Fig. 8. SEM image of the low angle impact zone at 50 ◦ C 500 mg/l solid loading conditions showing the carbide network protruding on the surface of cast Stellite 706.
Fig. 11. Anodic polarisation curves of HIPed Stellite 6, at 200 mg/l solid loading at 20 ◦ C and 50 ◦ C at 17 m/s in 3.5% NaCl solution.
descending order as follows: HIPed Stellite 706 > Cast Stellite 706 > HIPed Stellite 6 > Cast Stellite 6.
4. Discussion
Fig. 9. Secondary electron SEM image active wear zone (zone I) on the surface of HIPed Stellite 706 showing heavy plastic deformation at 50 ◦ C, 500 mg/l solid loading conditions.
The loss of material by erosion–corrosion is a complex process affected by many interacting variables. Thus changing any variable can have an important consequence on the behavioural pattern of the total system. In this paper two different cobalt-based super alloys, Stellite 6 and Stellite 706
Fig. 12. Effect of solid loading on the anodic polarisation of HIPed Stellite 706 at 20 ◦ C compared with the anodic polarisation curve under static conditions at 17 m/s in 3.5% NaCl solution.
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Table 7 Corrosion current densities (icorr ) values (A/cm2 ) at two different sand loading ant two different temperature Cast Stellite 6 20 ◦ C 50 ◦ C
HIPed Stellite 6
Cast Stellite 706
HIPed Stellite 706
200 mg/l
500 mg/l
200 mg/l
500 mg/l
200 mg/l
500 mg/l
200 mg/l
500 mg/l
7 12
13 33
4 8
10 18
3 8
10 19
3 6
8 16
produced by two different methods (casting and HIPing) have been subjected to liquid–solid erosion–corrosion where the impact velocity, erodent characteristics and impact angle remain constant. Erodent concentration and temperature of the slurry were changed. The effect of increase in temperature and increase in solid loading were discussed in detail in the previous work [5]. The difference in the erosion–corrosion resistance of the two cast and HIPed alloys are primarily due to the differences in: (a) microstructure – which affects mechanical integrity; (b) corrosion behaviour – which is important for the formation of the passive film and the extent of charge transfer during depassivation; (c) strength of matrix (affected by alloying elements) – which affects mechanical integrity and resistance to charge transfer. In terms of the microstructure effects, the cast samples of both materials experienced higher weight loss due to mechanical erosion compared to HIPed samples because they had larger matrix areas between the dendrites that were unprotected by the harder carbides. The coarser carbide size meant that, when carbide removal did not occur, large pieces of matrix were removed, which also accelerated the wear rate. Under erosion–corrosion conditions these unprotected matrix areas also plays an important role as more matrix area will be in contact with the corrosive environment. To understand why the alloy containing molybdenum has better erosion–corrosion resistance compared to the alloy that contains tungsten the three factors listed above as (a), (b) and (c) are further considered.
Fig. 13. Effect of solid loading on the anodic polarisation of cast Stellite 706 at 20 ◦ C compared with anodic polarisation in static conditions.
When the SEM images of Stellite 706 (both cast and HIPed) and Stellite 6 (both cast and HIPed) (Fig. 2) are compared it can be clearly seen that in addition to the formation of chromium-rich carbides in the alloys with molybdenum there is a formation of a molybdenum-rich carbide. The formation of these secondary carbides depends on the solubility limits of molybdenum and tungsten in cobalt. The solubility limit of both molybdenum and tungsten in cobalt extends to a maximum of 40% by wt. at the eutectic temperature [19]. As molybdenum (atomic weight = 94.94 g) has twice as many atoms as tungsten (atomic weight = 183.84 g) for the same weight percentage, there will be an excessive atomic concentration of molybdenum when a comparable wt.% is added. Because of this the propensity for molybdenum-rich carbides to form is higher. The main role of both tungsten and molybdenum in cobalt-based alloys is in solid-solution strengthening. In Stellite 6 as all the tungsten added takes part in solid solution strengthening, there is no excessive tungsten to form secondary tungsten-rich carbides and, as such, no tungsten-rich carbide formation is observed. As summarised in Table 6 carbide volume fractions on HIPed Stellite 706 are much higher compared to the HIPed Stellite 6. The volume fraction of carbide directly affects the erosion rate of the alloy and is one major contributory factor to Stellite 706 having superior erosion–corrosion resistance. As summarised in Table 3, the microhardness of cast Stellite 6 is higher compared to HIPed Stellite 6, but the erosion–corrosion resistance of the HIPed alloy is higher (Fig. 5). Again comparison of the both cast alloys of Stellite 6 and Stellite 706 (Fig. 5), the alloy with lower hardness value has higher erosion–corrosion resistance which shows that there is no direct relation between the hardness and erosion–corrosion resistance which is not unexpected if the controlling factors in erosion–corrosion are considered. In erosion alone little correlation between hardness and wear rate is seen partly due to the ideas of Hutchings and Levy [21] who postulated that the temperature rise during particle impact “anneals out” the effects of the hardening mechanism. Surface hardness after erosion might be a useful parameter to rank the erosion resistance of the materials. In addition in erosion–corrosion there is a substantial effect of corrosion which is unlikely to be directly linked to hardness and is more likely to be affected by depassivation/repassivation characteristics. Since the shape, size and the composition of the carbides are different in both cast and HIPed microstructure, support from the matrix is essential. In the case of abrasion wear it has been shown that sufficient support by the matrix can prevent
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Table 8 Normalised alloy content of Stellite 6 and Stellite 706 Alloy Cast Stellite6 HIPed Stellite 6 Cast Stellite 706 HIPed Stellite 706
Normalised alloy content (NAC) 1 1.2 1.36 1.39
pulling out of the carbides [23]. This is also the case where sand particles impact at low angles, as is the case in zone II in Fig. 6. The strength of the matrix is considered in detail here. Kosel et al. [24] showed the effect of the matrix strength can affect the wear rate. They used the normalised alloy content (NAC) as a measure of the matrix strength. They defined the normalised alloy content as the sum of the weight percentages of nickel, vanadium, tungsten; the alloys that are used as solid-solution strengtheners to give an approximate measure of the degree of solid-solution strengthening. By using the same approach the NAC of the alloys Stellite 706 and 6 were calculated and normalised against cast Stellite 6 as shown in Table 8. As shown in Fig. 14 there is linear relationship between the NAC and the erosion–corrosion resistance of the alloys. The alloys with high NAC gave a lower weight loss in all test conditions. This strain hardening of the matrix by alloy addition is explained through the change in the stacking fault energy (SFE). Excellent erosion resistance of cobalt-based Stellite alloy is greatly attributed to the low SFE. Molybdenum, tungsten and chromium tend to stabilize the hcp allotrope and decrease the SFE [18]. The lower the SFE (the greater the width of stacking fault) the more difficult cross-slip is and the higher the work-hardening and the strain to fracture. The materials with low SFE tend to strain harden rapidly, and show the highest erosion resistance. Both molybdenum and tungsten lowers the SFE, and so it is likely that in both Stellite 6 and 706 there will be some effect of SFE reduction. However the exact degree of this effect still needs further investigation. In terms of the corrosion behaviour under liquid–solid impingement, generation of fresh metal surface and the repassivation ability are two important factors particularly in the
Fig. 14. Total weight loss test results plotted vs. normalised alloy content.
case of discontinuous impacts of solid particles where depassivation caused by solid particle impacts is followed by a process of new surface repassivation. The features of anodic polarisation data can be very useful in determining the nature of the liquid–solid particle interactions with the solid surface and the associated corrosion activity. Both cast and HIPed Stellite 706 showed a pseudo passive region (region II in Fig. 12) in which the measured current did not change with the potential but where there is clear charge transfer and loss of true passivity. This indicates that the surface protective film is continuously destroyed by the mechanical factors but that the degree of charge transfer associated with the impacts is lower than on Stellite 6. Comparison of the corrosion current densities in Table 7 shows that in all conditions HIPed Stellite 706 is lowest which in turn means the resistance to breakdown of the passive film is highest. This can be due to the fact that it contains more carbides so that less matrix area is in contact with the corrosive liquid and the chemical composition of the matrix contains more corrosion-resistant elements like chromium to maintain integrity against charge transfer. Also comparison of the is (which is the current value at the start of the pseudo-passive region) as shown in Fig. 15, shows that the same trend in icorr is observed. A higher icorr (corrosion in the active state) also translates to a higher is (charge transfer associated with depassivation regions) such that the resistance of HIPed Stellite 706 is greatest and cast Stellite 6 lowest. Comparison of the two conditions (Fig. 15a
Fig. 15. Comparison of the anodic polarisation curves of Stellite 6 and Stellite 706 at (a) 18 ◦ C, 200 mg/l and (b) 50 ◦ C, 500 mg/l.
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and b) show that increasing the solid loading and temperature, increases is since more charge transfer occurs due to the impacts.
5. Conclusions The erosion–corrosion of cast and HIPed Stellite 706 has been examined in aggressive erosion–corrosion conditions and compared with cast and HIPed Stellite 6 alloy. The following conclusions are reached. All Stellite alloys for the conditions studied have superior resistance to superduplex (UNS S32760) and austenitic (UNS S31603) stainless steels. HIPed alloys show improved resistance to cast alloys but the degree of improvement is less in Stellite 706 than Stellite 6. 1. Microstructure plays a crucial role in the erosion– corrosion performance of the alloys. Addition of Mo is much more effective than W addition in conferring erosion–corrosion resistance to cobalt-based alloys due to formation of secondary carbides. 2. The strength of matrix, which is affected by alloying elements, is also to be considered as an important factor to affect the erosion–corrosion performance of the alloys. 3. The rate of the charge transfer during depassivation and the rate of repassivation of the surface is affected by the microstructure and alloying elements in the matrix. Acknowledgements The authors acknowledge the financial support to UM from Deloro Stellite and Heriot-Watt University. References [1] X. Hu, A. Neville, An examination of the electrochemical characteristics of two stainless steels (UNS S32654 and UNS S31603) under liquid–solid impingement, Wear 256 (2004) 537–544. [2] D.L. Klastrom, Wrought cobalt-based superalloys, J. Mater. Eng. Perform. 2 (1993) 523–530. [3] A. Neville, M. Reyes, T. Hodgkiess, A. Gledhill, Mechanisms of wear on a Co-based alloy in liquid–solid slurries, Wear 238 (2000) 138–150. [4] A. Neville, T. Hodgkiess, Characterisation of high-grade alloy behaviour in severe erosion–corrosion conditions, Wear 233–235 (1999) 596–607.
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