coororsion

1. Wear 255 (2003) 146–156
Corrosion and erosion damage mechanisms during erosion–corrosion
of WC–Co–Cr cermet coatings
V.A. de Souza∗, A. Neville
School of Engineering and Physical Sciences, Corrosion and Surface Engineering Research Group,
Heriot-Watt University-Edinburgh, Edinburgh EH14 4AS, UK
Abstract
Experiments were performed in order to evaluate the behaviour of WC–Co–Cr coatings applied by the Super D-GunTM (Trademark
of Praxair S.T. Technology) thermal spray process in erosion–corrosion environments. The experiments were performed in 3.5% NaCl
with different silica sand loadings (200 and 500 mg/l). The results were compared with UNS S31603 and UNS S32760 stainless steels.
Measurements of the total material loss showed that the coating applied by Super D-Gun presented higher resistance to erosion–corrosion
compared to both stainless steel materials as expected. The connection between the erosion–corrosion behaviour of the Super D-Gun coating
and the microstructure is assessed. Scanning electron microscopy (SEM) images enabled the different mechanisms of tribo-corrosion to
be understood. With cermet materials there has been much discussion relating to the role of corrosion in removing the hard phase particles
which can breach tribological performance. The corrosion rate (icorr) of the coating was determined under erosion conditions and an
evaluation of the influence of corrosion and synergistic processes on the erosion process was made using applied cathodic protection.
© 2003 Elsevier Science B.V. All rights reserved.
Keywords: Cermet; Coating; Corrosion; Erosion; Electrochemical
1. Introduction
Thermal spray coatings are being used in a diverse range
of engineering applications to extend component life by
retarding wear degradation. Since their inception, thermal
spray technologies have been the subject of some studies
[1–7] involving optimisation of process parameters to con-
sistently yield coatings with good bond strength, minimum
residual stresses and low porosity. The present situation is
therefore one in which high quality coatings are routinely
produced by industrial processes.
Thermally sprayed and solid cermets have been the sub-
ject of numerous dry erosion and abrasion studies [3–11]. In
a study using slurry impingement methodology, Hawthorne
et al. [12] investigated the dry erosion resistance of HVOF
(high velocity oxy-fuel) WC cermets with Co- or Ni-based
matrices with 50 ␮m alumina particles at 84 m/s. They found
that all coatings exhibited a ductile response to erosion with
WC–12Co having the best erosion resistance due to the high
carbide content. Analysis of the surface of the WC–Co coat-
ing after erosion showed evidence of cutting, platelet forma-
tion and occasional carbide particle removal. In comparing
∗ Corresponding author. Tel.: +44-131-449-5111x4737;
fax: +44-131-451-3129.
E-mail address: v.a.d.souza@hw.ac.uk (V.A. de Souza).
dry erosion with slurry impingement they found that the ex-
tent of damage was orders of magnitude greater in dry con-
ditions, mainly due to the much greater real particle impact
velocities. The ranking of the materials was found to be the
same at 90◦ and 20◦ angle of impingement. In their work
no consideration was given to corrosion effects in erosion
processes.
O’Quigley et al. in 1997 [13] used the abrasion test spec-
ified in ASTM Standard B 611-85 to rank the performance
of WC–Co cermets with variations in the carbide size and
the Co content. They produced guidelines to assist in the
selection of WC–Co cermets for abrasion resistance. It was
acknowledged in their work that the ranking was only valid
for the specific test conditions and that departure from these
conditions will probably lead to a rearrangement of the rank-
ing. As a rough guide, hardness of the material was only
useful at low hardness values where abrasion damage was
due to plastic deformation. Liao et al. [14] more recently
used a similar methodology to determine the important pa-
rameters for achieving high abrasion resistance of WC–Co
cermets. They concluded that cohesion between the carbide
particles and the matrix and the hardness were the two most
important factors. Improved cohesion at the interface was
due to the amorphous or nanocrystalline interaction phases
in the coating. Presence of M6C or M12C phases were found
to indicate a strong cohesion.
0043-1648/03/$ – see front matter © 2003 Elsevier Science B.V. All rights reserved.
doi:10.1016/S0043-1648(03)00210-2

2. V.A. de Souza, A. Neville / Wear 255 (2003) 146–156 147
Table 1
Different coating compositions used in the work by Wentzel and Allen
[15]
Grade Binder Composition (wt.%)
WC Ni Cr Co
C6 Co 94 0 0 6
V6 Ni 94 6 0 0
V7 Ni–Cr 94 5.4 0.6 0
P6 Ni–Cr–Co 94 3.0 0.6 2.4
With the increased use of cermet coatings and solid cer-
mets in applications where corrosion can play a part in the
degradation process, it is becoming increasingly important
to be able to assess the effects of the joint erosion and cor-
rosion processes. Work in this area has therefore increased
since the start of the 1990s. Wentzel and Allen [15] inves-
tigated the behaviour of tungsten carbide hard metals with
different binders subjected to erosion–corrosion (using a
slurry of silica and water with 7 mass% of solids at 6.5 m/s)
using gravimetric and potentiodynamic techniques. In this
work four compositions were used in order to evaluate the
binder influence on the material performance (Table 1). It
was concluded that no simple relationship exists between
the erosion–corrosion performance and any one material
property. The slurry erosion resistance was improved by
alloying of the single component Co and Ni matrices. No
improvement in erosion–corrosion resistance was reported
by increasing the passivation of the binder.
Toma et al. [16] in a recent study of six thermal spray coat-
ings found that addition of Cr to a Co matrix in WC–Co cer-
mets increased resistance to erosion–corrosion. Work by the
authors [17–19] on the pure corrosion resistance of WC–Co
and WC–Co–Cr coatings also showed that the Cr changed
the principal mechanism of corrosion. Toma et al. also re-
ported that replacing some of the WC with another hard
phase particle (Cr3C2) can add resistance to the material in
terms of erosion performance. To optimise corrosion perfor-
mance the study concluded that a move from CoCr to NiCr
binders was effective.
In this paper the electrochemical behaviour of WC–Co–Cr
thermal sprayed coating is reported in severe impingement
conditions. The effect of the impinging jet on the corrosion
response is discussed. Through application of cathodic pro-
tection (CP) the effects of erosion, corrosion and interac-
tions between erosion and corrosion can be determined. In
the paper the relative importance of corrosion and erosion
effects in the material loss are discussed.
2. Experimental methods
The coating studied in this work is a Super D-Gun thermal
sprayed coating of the WC–Co–Cr generic type.
Characterisation of the coating microstructure was carried
out using light and scanning electron microscopy (SEM).
The environmental SEM is equipped with a LaB6 gun,
and is capable of operating as a conventional high-vacuum
SEM, or under low vacuum in ESEM mode. The ESEM
is fully equipped with a range of secondary electron (SE)
and back-scattered electron (BSE) detectors with an energy
dispersive X-ray (EDX) chemical analysis facility. Correc-
tions for atomic number, absorption and fluorescence (ZAF)
are achieved through a virtual standard calibration routine.
The crystalline phases of the coating were identified using
X-ray diffraction (XRD) and measurements were made
on 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).
The impingement apparatus comprised a liquid–solid jet
generated using a recirculating rig as shown in Fig. 1 and
described elsewhere [19]. The rig comprised a dual nozzle
system. The velocity of the jet for this study was kept con-
stant at 17 m/s. The solid loading in the 3.5% NaCl fluid
was varied at two levels of 200 and 500 mg/l. The sand size
distribution is given in Table 2. The temperature of the liq-
uid was 18 ◦C. For all tests the angle of impingement was
90◦. The surface area of the samples exposed to the jet was
3.8 cm2.
Electrochemical analysis was used in conjunction with
weight-loss analysis to determine the total material loss
(TWL) and to isolate the contributions due to pure corrosion
(C) and pure erosion (E). The corrosion rate was measured
in situ using a three-electrode electrochemical cell compris-
ing a Ag/AgCl reference electrode connected by means of a
salt bridge and a platinum counter electrode. DC anodic po-
larisation tests (in static conditions or under the impinging
jet) involved scanning the potential of the working electrode
(the specimen under examination) from the free corrosion
potential (Ecorr) in the more noble (positive) direction at a
fixed rate of 0.25 mV/s. The potential was scanned in the
positive direction until the current flowing in the external
circuit between the working and counter electrodes reached
a value of 500 ␮A/cm2. The anodic polarisation tests were
started after 30 min exposure to static saline solution or the
impinging jet.
To measure changes in the corrosion rate as a func-
tion of solid loading on exposure period to the impinging
jet linear polarisation tests were conducted. In these tests
the potential of the working electrode (the sample under
erosion–corrosion) was shifted at a rate of 15 mV/min from
Table 2
Sand size distribution for the HST congleton silica sand used in
erosion–corrosion impingement tests
Sand size (mm) Percentage of total mass
<160 1
160 15
180 37
250 21
300 20
>425 7

3. 148 V.A. de Souza, A. Neville / Wear 255 (2003) 146–156
Fig. 1. (a) The rig configuration used in the experiments and (b) the electrochemical set up on the nozzle.
0.05 V negative to the free corrosion potential to 0.05 V
positive to the free corrosion potential. The applied poten-
tial is then a linear function of the current density in the
external cell and changes in the polarisation resistance (Rp)
can be calculated using the slope E/ I. In conjunction
with these measurements the free corrosion potential (Ecorr)
was also measured as a function of the solid loading.
In order to isolate the erosion component of material
loss, the weight loss was measured after exposure to the
impinging jet with applied CP which minimised the corro-
sion current on the sample. Calculations of the equilibrium
electrode potential for the Co/Co2+ reaction [20] confirmed
that application of −0.8 V (SCE) as the protection potential
was sufficiently near to the equilibrium electrode potential
to reduce corrosion effects to a negligible level. However,
the potential is not in the regime where hydrogen evolution
would be a potential complication at the cathode interface.
The CP was applied by potentiostatic means and involved
maintaining the potential of the working electrode at a con-
stant value of either −0.5 or −0.8 V against the saturated
calomel electrode (0.755 V Ag/AgCl).
Following all liquid–solid tests the surface was examined
using light and SEM to determine the extent of degradation
and to identify the material loss mechanisms.
3. Results
3.1. Coating characterisation
Fig. 2 shows an image of the coating microstructure where
the hard phase particles (light) are clear. The coating clearly
shows a lamellar structure produced by the layers built up
during the spraying process. The thickness of the coating is
250–300 ␮m. The hard phase particles are typically 1–2 ␮m
in size. The mean hardness of the coating is 1435 HV with
a standard deviation of 312 HV.
Table 3 shows the results of seven EDX measurements
taken over the surface of the coating. Each EDX measure-
ment was taken from a region of 500 ␮m × 500 ␮m. From
this it is clear that there is little spatial variation in the com-
position of the coating. The coating is nominally a WC–12%
Co–6% Cr. The oxygen content is approximately 1%.
The XRD trace for the coating is shown in Fig. 3. The
main constituents of the coating are identified. In compari-
son with WC–Co–Cr coatings sprayed using the HVOF pro-
cess, as reported in other work [17,18], it is evident that the
coating comprises a more complex microstructure. There is
evidence of decarburisation which has led to formation of
Fig. 2. Microstructure of the coating showing the hard phase particles
and the lamellar structure.

4. V.A. de Souza, A. Neville / Wear 255 (2003) 146–156 149
Table 3
EDX measurements on regions of 500 ␮m×500 ␮m on the coating surface
Element
(wt.%)
Sample number
1 2 3 4 5 6 7
C 2.5 2.6 3.1 3.5 3.7 3.5 3.5
O 0.9 0.8 1.0 0.1 0.7 1.1 0.9
Cr 5.8 5.7 5.6 5.9 6.0 5.8 5.9
Co 11.4 10.6 11.0 11.4 11.2 11.4 11.3
W 79.4 80.3 79.3 79.1 78.4 78.3 78.5
tungsten carbide forms other than simply the tungsten mono-
carbide species as well as ␩ phases. The amorphous phases
identified by the broad peaks in the 2θ range from 35◦ to
45◦ represent the dissolution of Co to form Co3W3O [7].
However, a residual amount of Co as metal is present which
was confirmed by XPS by Souza and Neville [21]. Also the
broad peaks could be accounted for by the (W, Cr)2C and
WC(1−x) phases.
3.2. Corrosion characteristics
3.2.1. Static conditions
The Ecorr of the coating in static 3.5% NaCl, measured
prior to starting the anodic polarisation curve, was consistent
at 285 ± 15 mV.
Fig. 4 shows the anodic polarisation curve for the
WC–CoCr coating in static 3.5% NaCl at 18 ◦C. For com-
parison the high grade superduplex stainless steel and the
UNS S31603 austenitic stainless steel anodic polarisation
curves under identical conditions are presented. It can be
seen that there are some similarities to the stainless steel
response and some contrasting features. The most important
contrasting feature is the rise in current at the potential near
to Ecorr on the coating compared with the extremely low
values of current (<1 ␮A/cm2) exhibited by both stainless
Fig. 3. XRD trace from the as-received coating.
Fig. 4. Anodic polarisation curve measured on WC–Co–Cr coating in
static conditions compared with UNS S32760 (superduplex) and UNS
S31603 (austenitic) stainless steels.
steels. This is indicative of some corrosion activity existing
on the coating compared with the truly passive response of
both stainless steels. The breakdown potential exhibited on
UNS S32760 and UNS S31603 at 1 and 0.3 V, respectively,
is characteristic of passivity breakdown and represents the
potential at which localised corrosion is forced to occur. On
the coating it is clear that two such potentials are exhibited
at 0.55 and 0.93 V and these are related to localised corro-
sion at the carbide/binder interfaces and pitting of the CoCr
binder, respectively, as reported in another communication
[17]. From the anodic polarisation curve, it is possible to
determine a dissolution rate for the material through ex-
trapolation of the E–log I curve when plotted in the region
near to (from 50 mV to not more than 200 mV) Ecorr. This

5. 150 V.A. de Souza, A. Neville / Wear 255 (2003) 146–156
Fig. 5. Localised corrosion attack around the interface between the hard
phase particles and the matrix after anodic polarisation in static 3.5%
NaCl at 18 ◦C.
enables the corrosion current density to be determined and
for the coating in static conditions this is 1.3 ␮A/cm2. Whilst
this is an extremely low value of corrosion current the im-
portant aspect to remember is that the coating comprises a
(W, Cr)2C and ␤-WC(1−x) (which will have extremely low
electrochemical activity) and corrosion proceeds primarily
by dissolution of the Co phase in the first instance. This was
confirmed by ICP in [21]. The exposed area of the matrix is
much less than the ceramic phase (see Fig. 2) and, as such,
these corrosion currents (calculated from the total exposed
cermet surface) give underestimated dissolution rates (in
terms of thickness) for the metal phase. This is evident from
Fig. 5 where it is clear that severe localised dissolution of
the binder around the hard phase particle/matrix interface
has occurred. There are obvious implications for the role
of corrosion in erosion–corrosion when the integrity of the
bond can be lost through binder dissolution.
Measurement of the linear polarisation characteristics of
the WC–Co–Cr coating in static conditions produced the
E/ I value for Rp of 3500 . By assuming the values for
βa, βc of 0.1 V in Eq. (1) for calculation of the corrosion
Fig. 6. The free corrosion potential (Ecorr) as a function of solid loading under impingement conditions at 17 m/s and at normal impact angle. Solution
is 3.5% NaCl at 18 ◦C.
rate, the icorr value can be estimated. The resulting icorr value
is calculated
E
I
=
βaβc
2.3icorr(βa + βc)
= Rp (1)
to be 6.08 ␮A which, for the specimen area of 3.8 cm2 is a
corrosion current density of 1.6 ␮A/cm2. This value is very
close to the value determined by anodic polarisation in static
3.5% NaCl.
3.2.2. Corrosion characteristics under impingement free
corrosion potential
The free corrosion potential as a function of solid loading
is shown in Fig. 6. On subjecting the sample to impingement
(with no solids) there is a shift in the positive direction of
15 mV from the value recorded in static conditions. There
is then only a very small shift of <10 mV from the value at
zero solids (with impingement of liquid only) to the values
at 100 mg/l. This is in contrast to the response of stainless
steels under impingement as will be discussed later in the
paper. At between 100 and 200 mg/l there is the largest shift
in the active direction and beyond that solid loading only
very small active shifts are recorded.
3.2.3. Anodic polarisation
Fig. 7 shows the forward anodic polarisation curves for
the WC–Co–Cr coating in static conditions and under im-
pingement at 200 and 500 mg/l solids. As stated previously
in static conditions there are very low currents in the region
near to Ecorr and this is manifested in a very low icorr. Un-
der impingement at 200 mg/l the currents are larger over the
entire potential range. In the region near to Ecorr the cor-
rosion current densities can be determined and from Tafel
extrapolation the values in Table 4 are determined. This in-
dicates that there is a large effect of the impingement on
the corrosion current density—depolarisation of the anodic
dissolution reaction.

6. V.A. de Souza, A. Neville / Wear 255 (2003) 146–156 151
Fig. 7. Anodic polarisation curves under impingement conditions for the
WC–Co–Cr coating in 3.5% NaCl at 17 m/s at 18 ◦C.
Table 4
Corrosion current densities measured as a function of solid loading under
impingement
Condition Corrosion current
density, icorr (␮A/cm2)
Static 1.3
Liquid–solid impingement, 200 mg/l solids 5.8
Liquid–solid impingement, 500 mg/l solids 21.8
3.2.4. Linear polarisation
Fig. 8 shows the values of Rp ( E/ I) determined for
the five solid loadings from 0 to 500 mg/l. In comparing
the value for Rp between static conditions and impingement
with no solids there is no significant difference. However,
on addition of 100 mg/l solids to the impinging stream there
is a significant decrease in the Rp value. The largest changes
of Rp occur at the solid loadings up to 200 mg/l after which
the Rp becomes much less dependent on solid loading.
Fig. 8. Polarisation resistance (Rp) calculated from the linear polarisation measurements across a potential range of ±50 mV from the free corrosion
potential over a range of solid loadings from 0 to 500 mg/l.
3.2.5. Cathodic polarisation
Depolarisation of the cathodic reaction is clear as the
WC–Co–Cr coating is subjected to erosion–corrosion as
shown in Fig. 9. In static conditions a small region between
−0.8 and −1 V exists where the current rises much less as a
function of potential signifying some partial diffusion con-
trol. The addition of solids to the slurry stream (200 and
500 mg/l) increases the cathodic current density measured
at each potential compared to the static condition. The sig-
nificance of these curves is in the determination of the opti-
mum potential for application of CP. This will be discussed
in great detail in Section 4.
3.3. Erosion–corrosion material loss
3.3.1. Volume loss
Fig. 10 shows the total volume loss recorded for the coat-
ing and the two stainless steels at two solid loadings, 200
and 500 mg/l, respectively. As expected the stainless steels
have a lower resistance to erosion–corrosion than the cer-
met coating. The coating exhibits an enhanced performance
margin over the materials UNS S31603 and UNS S32760
as the solid loading increases. At 200 mg/l solid loading the
coating exhibited a mass loss of 64 and 76% of the UNS
S31603 and UNS S32760 value, respectively, at 500 mg/l
the figures were lower at 55 and 68%, respectively, thus in-
creasing the margin.
3.3.2. Material loss mechanisms
Fig. 11 shows the surface of the WC–Co–Cr coating
after erosion–corrosion for 8 h at two solid loadings and
three regions on the surface are clearly defined on each
surface. Region 1 is the central zone where the surface is
dulled considerably due to the erosion–corrosion damage.
This ‘dulling’ of the surface is due to the development of
craters which appear as macro pits on the surface. These are

7. 152 V.A. de Souza, A. Neville / Wear 255 (2003) 146–156
Fig. 9. Cathodic polarisation curves measured in static and liquid–solid impingement conditions.
Fig. 10. Total volume loss measured for the coating and the two stainless steels under liquid–solid impingement.
Fig. 11. Wear scar on the surface of WC–Co–Cr after exposure to the
impingement with overall surface image with three different regions.
evident on surfaces without and with CP applied and hence
are a mechanical feature of the damage. Fig. 12a shows the
boundary between the central dull region and the concen-
tric ring described as the erosion ‘halo’ (Region 2) in the
literature [22,23]. Inside Region 1 the damage reflects the
fact that impacts are nominally at high angle and it can be
seen that there is extensive brittle fracture and removal of
the hard phase particles (Fig. 12b). Region 3 is the outer af-
fected zone where damage due to impingement is still clear
but there are clear differences when compared to Region 1.
Fig. 12c shows the surface of Region 3 and it is clear that
there is directionality associated with the damage. There
is evidence of material (matrix) flow along the direction of
the flow out from the central zone. There is far less evi-
dence of hard particle removal than in Region 1 although
there are still sites where macro-cracking is observable.

8. V.A. de Souza, A. Neville / Wear 255 (2003) 146–156 153
Fig. 12. Damage in: (a) Region 1; (b) Region 2; (c) Region 3; (d) shows the form of the wear scar profile typically found on eroded surfaces of
WC–Co–Cr coating when exposed to liquid–solid impingement at 500 mg/l solids for 8 h.
Fig. 13. CP currents for WC–Co–Cr under liquid–solid impingement for 200 mg/l solids at −0.5 and −0.8 V (SCE).
Fig. 12d shows the profile of the wear scar determined from
profilometry.
3.3.3. Application of CP
Fig. 13 shows the CP currents as a function of time with
200 mg/l solids over the 8 h test period at two potentials:
−0.5 and −0.8 V. The magnitude of the current is increased
as the overpotential for the applied CP potential increases.
Also, as the solid loading increases the amount of current
required to maintain the CP potential increases as illustrated
by the total charge during CP figures in Table 5.
Table 5
Total charge during CP under impingement for 8 h period
Applied CP potential
(V) (SCE)
Total charge during CP (C)
200 mg/l solids 500 mg/l solids
−0.5 90.4 139.1
−0.8 646 812.8

9. 154 V.A. de Souza, A. Neville / Wear 255 (2003) 146–156
Fig. 14. Weight losses for WC–Co–Cr for liquid–solid impingement without and with applied CP.
Fig. 14 shows the reduction in material loss measured in
the presence of applied CP compared with the material loss
measured with no applied CP. It can be seen that there is
a significant effect at both solid loadings of applying CP at
the level of −0.8 V. At −0.5 V the effect is slightly more
complicated and will be discussed later.
4. Discussion
The use of thermal spray coatings for reducing material
degradation in tribo-corrosion environments has increased
in recent years and successful applications in pump compo-
nents, valves and other equipment subjected to impinging
solid-laden stream are commonplace. However, there is still
a need to define the limitations for WC–Co–Cr and other
cermet coatings, especially where corrosion and joint ero-
sion/corrosion mechanisms can affect performance. It has
been demonstrated in the previous work by the authors that
localised corrosion attack at carbide/matrix interfaces can
lead to loss of support for carbides and also dissolution of
carbides when the local environment becomes aggressive
during localised corrosion. In this current paper the em-
phasis is on investigating how corrosion and erosion inter-
act and the test methodology has been devised with this in
mind.
4.1. Overall performance
WC–Co–Cr from the thermal spray process investigated
in this study showed superior resistance to erosion–corrosion
than both of the stainless steels (UNS S31603 and
UNS S32760) as expected. As the conditions of the
erosion–corrosion became more severe (increased solid
loading) the difference between the cermet and stainless
steels was enhanced. Hawthorne et al. [12] found ratios of
UNS S31603/WC–Co–Cr of almost 6:1 in slurry impinge-
ment tests which is much greater than those experienced
here but it is important to note that the slurry used in their
tests was not corrosive and the solid loadings were 9.1 wt.%
compared with the 500 mg/l (0.05%) in this study and
also the difference in composition between the D-Gun and
HVOF coatings can account for the difference in wear rates.
It is apparent from the nature of the wear scar on
WC–Co–Cr (showing highest degradation at a nominal 90◦
angle) in contrast to the stainless steels, as reported else-
where [23] that the material loss mechanisms are somewhat
different and some analogies can be made from considera-
tion of ductile and brittle material behaviour as was defined
in dry erosion in the early work of Finnie [24]. Considering
the transport of particles in this dilute liquid–solid stream, it
is apparent that at the centre of the jet (Region 1 in Fig. 11)
the particles will nominally impact at 90◦. On moving out
from the centre of the jet the surface will be exposed to
impacts at a lower angle and due to squeeze film effects the
particles may not indeed impact at all. If the mechanical
degradation laws relationships for dry erosion, established
first by Finnie [24] are transferred across to liquid–solid
impingement it may be assumed that for a ductile material
the highest loss of depth will occur out from the centre
(normally at an angle of around 30◦) and for stainless steels
the “w” shape confirms this to be the case. Also, for the
more brittle WC–Co–Cr the material loss is highest at 90◦.
The “u” shape of the wear scar confirms this to be the case.
4.2. Corrosion/erosion interactions
Application of CP at −0.8 V (SCE) was effective at both
solid loadings of reducing the corrosion component of the
damage to a negligible level as was proved to be the case
in [22]. At this level of CP the erosion damage is isolated
and it can be seen that at 200 and 500 mg/l the damage is
very much dominated by erosion processes. For 200 mg/l
E/TWL is 0.836 and at 500 mg/l it is 0.877. Hence as the

10. V.A. de Souza, A. Neville / Wear 255 (2003) 146–156 155
solids level increases the damage becomes more and more
dominated by erosion processes. This was in contrast to the
situation in solids-free impingement of an HVOF WC–CoCr
coating where the damage was dominated by corrosion
and corrosion-related (synergistic) processes [20]. Erosion,
playing a large part of the damage in slurry impingement
is often the case in aggressive conditions and it has been
shown to be the same on drill bit material (Stellite X-40)
[19] but on austenitic stainless steels the opposite was the
case [23] and corrosion became a higher proportion of the
damage as the solid loading increased. This was apparently
due to the crucial repassivation events which occur on solid
impact on stainless steels and the corrosion damage is criti-
cally affected by the ability to repassivate once damage has
occurred.
In terms of corrosion it is interesting to note that the
effect of erosion processes on corrosion of the cermet is
significant and that corrosion rates are increased by a fac-
tor of 4.46 at 200 mg/l solids and 16.8 at 500 mg/l solids
compared to static corrosion rates. It is known from XPS
that Cr in the coating forms oxide (Cr2O3) which will pre-
vent substantial charge transfer [21]. It is therefore likely
that under E/C the oxide on the metal phase is removed
and corrosion (at a rate greater than in static conditions)
can proceed. Hence impingement by impacting solids will
have the effect of removing any passive film on the binder
and render the material active and enable charge transfer.
Hence the large factor of Ce is in line with what would be
expected for this type of binder. Table 6 shows the values
of the components of material loss (TWL) according to
TWL = E + C + Ce + Ec (2)
where E is the pure erosion material loss determined through
application of CP, C and Ce are the static corrosion com-
ponent and the effect of erosion on corrosion, respectively,
and Ec the effect of corrosion on erosion (often referred
to as the synergy).
Corrosion measurements under erosion–corrosion condi-
tions have been conducted in this study using polarisation
techniques and then Tafel extrapolation to evaluate the cor-
rosion current density at the free corrosion potential. The
corrosion current densities established from these measure-
ments were used to evaluate the components of corrosion
damage in Table 6. As part of this investigation Rp mea-
surements were made at various solid loadings and it was
shown that the Rp decreased as the solid loading increased.
The trend was in line with the increase icorr measured
during impingement at 200 and 500 mg/l solids. However,
Table 6
Components of material loss (mg) attributed to different degradation
processes on WC–Co–Cr under liquid–solid impingement in 3.5% NaCl
E C Ce Ec
200 mg/l 5.1 0.04 0.15 0.81
500 mg/l 11.5 0.04 0.67 0.89
there is an important consideration to be made in that the
ratio of icorr at the two solid loadings does not match the
ratio of 1/Rp values at the two solid loadings. This therefore
indicates that for Rp measurements to give a good approx-
imation to the corrosion rates then the values of βa and βc
must be established under erosion–corrosion conditions and
these cannot be routinely assumed to be constant. Otherwise
erroneously low changes in corrosion current as a function
of solid loading would be estimated.
Interestingly at −0.5 V applied CP the damage was re-
duced at the lower solid loading of 200 mg/l to a comparable
value at −0.8 V but at the higher solid loading the applied
potential of −0.5 V was not effective in reducing the cor-
rosion component of damage. Two links to electrochemical
measurements can be made in this respect. Firstly the free
corrosion potential at 500 mg/l is more negative than at
200 mg/l and hence the CP overpotential is reduced and as
such the CP efficiency would be expected to drop. Also, as
the solid loading increases the corrosion current increases
and so the amount of current required to protect the surface
increases and cannot be sustained at −0.5 V. Hence it can be
deduced that for erosion–corrosion the choice of potential is
vital.
5. Conclusions
The study has demonstrated the following:
• WC–CoCr thermal sprayed coatings can provide good
protection against wear and corrosion in liquid–solid im-
pingement when compared with stainless steels.
• The mechanisms of damage are dominated by erosion
processes but corrosion is affected by erosion processes
and is more important at the lower solid levels.
• Monitoring of corrosion rates requires care to be taken in
interpreting linear polarisation data.
• CP can offer protection from erosion–corrosion damage—
the extent of the protection depends on the erosion–
corrosion severity.
Acknowledgements
The authors acknowledge Weir Pumps Ltd., Glasgow and
Greenhey Engineering Services, England, for the financial
support provided to V.A. de Souza.
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