Project report helping Notes of Sulrry Erosion testing Machine.

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Slurry Erosion Mechanism of Hydroturbine Steel: Effect of Operating
Parameters
Article in Tribology Letters · November 2013
DOI: 10.1007/s11249-013-0213-z
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2. ORIGINAL PAPER
Slurry Erosion Mechanism of Hydroturbine Steel:
Effect of Operating Parameters
H. S. Grewal • Anupam Agrawal • H. Singh
Received: 10 May 2013 / Accepted: 2 September 2013 / Published online: 13 September 2013
Springer Science+Business Media New York 2013
Abstract Performance of hydropower plant is severely
affected by the presence of sand particles in river water.
Degree of degradation significantly depends on the level of
operating parameters (velocity, impingement angle, con-
centration, particle size and shape), which is further related
to erosion mechanism. In this investigation, the effect of
some of these operating parameters on erosion mechanism
of generally used hydroturbine steel, CA6NM (13Cr4Ni),
is reported. Morphology and variation in the martensite and
austenite phases of the eroded surfaces were investigated
using SEM and XRD. It was observed that velocity and
impingement angle affect the erosion mechanism of
CA6NM steel. Erosion mechanism was also significantly
affected by the radial distance from the impact zone. Pri-
mary mechanism responsible for the removal of material at
normal impingement angle was the formation and removal
of platelets. At acute impingement angle, ploughing was
observed to be one of the prime mechanisms responsible
for the loss of the material. Other than these two well-
known erosion mechanisms, the presence of another two
erosion mechanisms was also observed. Models have been
proposed for these unfamiliar erosion mechanisms. Inter-
action amongst different operating parameters was studied
using line and contour plots. It was observed that the
interaction between velocity and concentration was most
significant. Using the experimental results, a statistical
model based on regression approach was developed.
Validity of this statistical model was checked using the
experimental results from the literature and present study.
Keywords Erosive wear Wear mechanisms
Power generation Steel
1 Introduction
Degradation of surfaces upon interaction with solid parti-
cles entrained in a liquid medium is termed as slurry ero-
sion. The impacts caused by these solid particles result in
the generation of stresses of a highly complex nature in the
target material. The damaged surface becomes prone to
further deterioration through promotion of cavitation and
turbulence in the flow field. With the development of tur-
bulence, the probability of multiple impacts by particles
also increases [1]. Slurry erosion is affected by numerous
parameters such as velocity of the particles, impingement
angles, particle flux, shape, size and morphology of parti-
cles, mechanical properties of both target as well as par-
ticles, and level and quantity of defects in particles and
target material.
Slurry erosion is the major cause responsible for the
degradation in fluid machineries, other than cavitation
erosion. Hydroturbines, especially those located on rivers
originating from Himalayas, are severely affected by slurry
erosion. It has been estimated that India alone faces a loss
of US $ 120–150 million as a consequence of slurry ero-
sion in hydroturbines [2]. In China, turbines installed in
Three Gorges Dam on Yellow River have been reported to
suffer high degree of damage due to slurry erosion.
Hydropower plants in Nepal also face challenges while
operating in highly concentrated sandy water [3]. The
problem is so severe that some of these plants need to be
shut down during monsoon season due to a considerable
increase in the concentration of sand. Normally, plants
continue to work with sand laden water up to a certain
H. S. Grewal A. Agrawal H. Singh ()
School of Mechanical, Materials and Energy Engineering, Indian
Institute of Technology Ropar, Nangal Road, Rupnagar 140001,
Punjab, India
e-mail: harpreetsingh@iitrpr.ac.in; hn1998@gmail.com
123
Tribol Lett (2013) 52:287–303
DOI 10.1007/s11249-013-0213-z

3. threshold value of concentration, usually 5,000 ppm
(0.5 wt% of sand). Beyond this threshold value, plants are
normally shut down to avoid significant damage to the
submerged parts. Some severally affected parts of hydro-
turbines are impellers, guide vanes, buckets, nozzles,
spears and labyrinth seal.
Commonly used structural material for hydroturbines is
CA6NM (ASTM 743). Other grades such as 16Cr5Ni,
CF8M and 13Cr1Ni have also been suggested; however,
CA6NM (13/4) remains a preferred choice, owing to its
good resistance against corrosion and cavitation erosion,
high impact and fracture toughness [4]. 13/4 steel is also
finding its use in other fluid machineries and components
such as pumps, compressors and valves [5–7]. Therefore,
it is vital to understand the performance of this steel in
slurry erosion condition. Although a good wealth of lit-
erature is available on the topic, slurry erosion, yet studies
related to 13/4 steel are quite limited in number. In most
of these studies, 13/4 steel has been used as a reference
material for evaluating the performance of different sur-
face modifications techniques against slurry erosion. For
improving the performance of this steel, it is required to
understand how different operating parameters affect the
erosion mechanism. Understanding the mechanism of
material removal is important for improving the working
life of the components and reducing the downtime of the
plants.
Detailed survey of the literature related to slurry erosion
performance of 13/4 steel revealed some information
regarding the erosion mechanism of this steel. Table 1
summarizes results from the literature related to the
experimental conditions used for slurry erosion testing of
13/4 steel. Chauhan et al. [8] studied the erosion response
of 13/4 steel against nitronic steel using solid particle
erosion test rig at both 30 and 90 impingement angles.
The authors observed a constant erosion rate during testing
for a period of 120 min, which was explained on the basis
of strain hardening capability of this steel. Ploughing was
observed to be the operative mechanism responsible for the
removal of material at an angle of 30, whereas at 90,
formation of deep craters was the cause for removal of
material. Higher mass loss was observed at 90 impact
angle in comparison with 30 impact angle. Grewal et al.
[9] conducted slurry erosion tests on 13/4 steel at different
sets of velocity, impingement angles, concentrations and
particle sizes. It was concluded that maximum erosion
occurs at 90 impingement angle. Erosion rates increased
linearly with the corresponding increase in velocity and
concentration.
Manisekaran et al. [10] evaluated the performance of
laser and pulsed plasma surface-modified and unmodified
13/4 steel. It was observed that maximum mass loss of 13/4
steel takes place at 60 impingement angle. It was also
observed that at 90 angle, the material removal took place
in the form of platelets, whereas at 45 angle, scars having
resemblance to low-angle abrasion were noticed. However,
no detailed investigation of effect of operating conditions
on erosion mechanism was undertaken. Shivamurthy et al.
[11] have also reported the similar kind of results under
identical testing conditions. It was noticed that the differ-
ence in the erosion rates at 60 and 90 angles was not
significant in both the preceding studies. The effect of
velocity in the range of 8–20 m/s on erosion rates was also
reported by Shivamurthy et al. [11]. Shivamurthy et al. [12]
further evaluated the effect of particle size on the erosion
behaviour of 13/4 steel. It was observed that with an
increase in particle size from 100 to 375 lm, the
impingement angle at which the maximum erosion takes
place reduces to 45 from 60.
Sugiyama et al. [13] observed that as the impact velocity
is increased, the impingement angle at which maximum
erosion takes place changes drastically for this steel. They
observed that at 10 m/s velocity, the maximum erosion rate
occurred at 60 impact angle. However, when testing was
conducted at 40 m/s impact velocity, erosion rates were not
affected by the impingement angle beyond 60. With an
increase in impingement angle beyond 60, the variation in
erosion rates with respect to impact angle was negligible.
Shivamurthy et al. [11] also reported stabilization in the
erosion rates beyond 60 impingement angle; however, the
impact velocity was only 12 m/s.
Romo et al. [14] also investigated the slurry erosion
behaviour of 13/4 steel against stellite weld overlay for the
range of impingement angles at a velocity of 25 m/s. It was
reported that for 13/4 steel, maximum mass loss takes place
at 30 and 45 impingement angles. Microploughing
combined with microcutting was observed as primary
erosion mechanism at these angles. At normal impinge-
ment angle, platelet mechanism played a dominant role
along with microploughing and microcutting. Numerical
predictions by Sugiyama [15, 16] also showed that maxi-
mum erosion for this steel occurs at 30 to 40 impinge-
ment angle.
Mann [17] evaluated the slurry erosion performance of
13/4 steel using a rotating disc-type apparatus. Operating
parameters were kept constant during the experimentation.
13/4 steel was used as a reference material for investigating
the performance of different surface modification tech-
niques, namely boronizing, plasma nitriding, plating, and
thermal spray coating. Chattopadhay [18] also studied the
slurry erosion performance of 13/4 steel with an aim of
quantifying its response against other test materials. Santa
et al. [19, 20] evaluated the erosion behaviour of 13/4 steel
in comparison with different protective coatings. Microp-
loughing and microcutting were identified as prominent
mechanism responsible for the removal of material.
288 Tribol Lett (2013) 52:287–303
123

4. Table
1
Summary
of
the
test
conditions
used
in
the
literature
for
erosion
testing
of
CA6NM
(13Cr4Ni)
steel
Reference
Type
of
test
rig
Test
conditions
Remarks
Erodent
Erodent
size
(lm)
Concentration
(kg/m
3
)
Impact
velocity
(m/s)
Impact
angle
()
Chauhan
et
al.
[8]
Jet
type
SiC
500–700
0.013
120
30
and
90
Testing
was
conducted
with
air
as
the
carrier
fluid,
which
obviously
do
not
represents
the
conditions
in
hydroturbines
Grewal
et
al.
[9]
Jet
type
Sand
5–20
20–60
30,
60
and
90
The
test
rig
was
recirculating
type,
which
could
result
in
the
slurry
ageing
Manisekaran
et
al.
[10]
Jet
type
SiO
2
150,
150–300
10
12
30,
45,
60
and
90
Shivamurthy
et
al.
[11,
12]
Jet
type
SiO
2
100
and
375
10
8–12
30,
45,
60
and
90
Sugiyama
et
al.
[13]
Jet
type
SiO
2
80
10
10
15
to
90
Romo
et
al.
[14]
Jet
type
SiO
2
212–300
*1
25
15
to
90
Testing
was
conducted
with
fresh
sand
each
time
for
5
min
Mann
[17]
Pot
type
Sand
135
2
75
NA
Fresh
sand
was
feed
continuously
Chattopadhay
[18]
Pot
type
Sand
Not
reported
Not
reported
20–22
NA
It
is
a
recirculating
type
test
rig.
Samples
remain
in
contact
with
slurry
for
only
one
half
of
the
rotation
Santa
et
al.
[19]
Modified
centrifugal
pump
SiO
2
212–300
100
5.5
and
10
NA
The
test
rig
was
recirculating
type,
which
could
result
in
the
slurry
ageing
Santa
et
al.
[20]
Jet
type
SiO
2
212–300
4
75
15
Testing
was
conducted
with
fresh
sand
for
4
cycles
of
1
min
each
Tribol Lett (2013) 52:287–303 289
123

5. As discussed above, a number of researchers have
evaluated the slurry erosion performance of 13/4 steel;
however, no systematic in-depth study has been conducted.
Although the effect of different operating conditions such
as velocity, impingement angle, concentration and particle
size has been studied, no detailed investigation was found,
which simultaneously aims at understanding the effect of
these operating parameters on the slurry erosion mecha-
nism of 13/4 steel. Material can perform differently under
different sets of testing conditions, which further depends
upon how the particles interact with the target surface.
Therefore, variation in operating condition could greatly
affect the erosion mechanism. Hence, it is important to
understand how 13/4 steel behaves under different sets of
experimental conditions and how these conditions affect
the erosion mechanism involved. Understanding the effect
of these operating conditions upon the erosion mechanism
could readily facilitate in improving the performance of
this steel and would enhance the knowledge related to the
erosion process.
2 Experimentation
2.1 Materials
In the present investigation, commonly used hydroturbine
steel, namely CA6NM (ASTM 743), was utilized. The
detailed composition of this steel is given in Table 2. As-
cast steel test block was machined, and samples of
20 9 20 9 5 mm were prepared using wire-cut electric
discharge machine (EDM). Prior to testing, samples were
polished using SiC emery papers down to 1,500 grit. This
was followed by polishing using 1-lm alumina powder
slurry on a disc polishing machine.
Sand used in present investigation was collected from
Sutlej River so as to perform testing under actual working
conditions. The sand collected was sieved using sieve
shaking machine, and all the sand particles having size
[300 lm were discarded. Particle size distribution of the
sand particles used for investigation is shown in Fig. 1a.
The analysis of the sand particles was undertaken using
scanning electron microscope (SEM) (Make: Jeol, Model:
6610 LV). SEM micrograph of the sand particles is shown
in Fig. 1b. The average size of sand particles is around
197 lm. The sand consists of irregularly shaped particles.
The circularity of the particles was found to be around 0.4.
The X-ray diffraction (XRD) (Make: PANalytical, Model:
X’pert Pro) analysis of the sand particles shown in Fig. 2
indicates silica (SiO2) as the major phase present in the
sand along with alumina (Al2O3) and ferrous oxide (Fe2O3)
as secondary phases. The microhardness of the sand mea-
sured using microhardness tester (Make: Wilson, Model:
MVH 402) was found to be around 1600 HV with a stan-
dard deviation of 200 HV.
2.2 Slurry Erosion Test Rig
For slurry erosion testing, jet-type test rig used in the
present investigation was designed and developed in-house.
The test rig falls under the category of non-recirculating
type, wherein sand particle after flowing through the system
once is not re-used. This helps in better simulation of actual
working environment. A systematic layout of the test rig
could be found in Fig. 3. In this rig, velocity, impact angle
Table 2 Composition of CA6NM steel used in experimentation
Element C Si Mn Ni Cr Mo W
Wt% 0.2 0.9 0.72 3.72 13.5 0.73 0.11
0
20
40
60
80
100
Cumulative
wt.
%
53-75
53 106-150
75-106
5 150-212
Particle siz
ze (µm)
212-300
(a)
(b)
Fig. 1 a Particle size distribution of the sand used for erosion testing
in the current investigation. b SEM image of the sand used for erosion
testing in the current investigation
290 Tribol Lett (2013) 52:287–303
123

6. of the specimen with reference to jet of slurry, concentration
and size of abrasive particles could be controlled indepen-
dently. The inter-dependency between the velocity and the
concentration was avoided through an improvement in the
design of the test rig as discussed in detail in previous
publication [21]. Slurry concentration was controlled by
feeding a known amount of sand into the mixer shown in
Fig. 3. Concentration of sand was calibrated by measuring
the amount of sand coming out of the nozzle. Nozzle of
4 mm diameter was made from tungsten carbide insert to
minimize the variation in nozzle diameter during experi-
mentation. Moreover, the outlet diameter of the nozzle was
measured periodically during experimentation to identify
any change. The flow/velocity of the slurry was monitored
continuously during experimentation using ultrasonic flow
meter also calibrated by measuring the quantity of water
coming out of the nozzle using discharge method.
2.3 Experimental Procedure
Erosion testing was conducted using slurry composed of
sand mixed in tap water. During experiments, three of the
operating parameters, namely velocity, impingement angle
and concentration, were varied using full factorial approach
as shown in Table 3. The nomenclature used for the
identification of samples could also be found in Table 3.
The first two digits of the nomenclature represent the
impact angle at which the experiment was conducted. Next
two digits indicate the test velocity, and last digit repre-
sents the concentration. Testing was conducted with fixed
stand-off distance of 20 mm. Experimentation conducted
using full factorial approach allows detailed investigation
of the interactions amongst the factors. The test was
interrupted regularly after every 60 s for mass loss mea-
surements and SEM analysis. The weight loss was mea-
sured using a precision weighing balance of 0.01 mg
accuracy. Prior to weight measurements, samples were
washed with acetone and dried in air to remove any
unwanted dirt or moisture content. Testing was continued
for 10 min with two samples tested under identical
experimental conditions to ensure repeatability. A regres-
sion-based model was derived using the experimental test
results. The predicted erosion rates were compared with
those obtained experimentally in the present study and
from the literature.
Fig. 2 XRD analysis of the sand used for slurry erosion experiments
.
Slurry Pump
Sand feeder
Re-circulated water line
Filter unit
Slurry containing line
Storage Tanks
Sample holder and
nozzle unit
Fig. 3 Schematic diagram illustrating slurry erosion test rig used for experimentation
Table 3 Experimental conditions used for slurry erosion testing of
CA6NM steel
Nomenclature Velocity
(m/s)
Impingement angle
(deg.)
Concentration
(wt%)
30041 4 30 0.1
30045 4 30 0.5
90041 4 90 0.1
90045 4 90 0.5
30161 16 30 0.1
30165 16 30 0.5
90161 16 90 0.1
90165 16 90 0.5
Tribol Lett (2013) 52:287–303 291
123

7. 3 Results and Discussion
3.1 Slurry Erosion
3.1.1 Effect of Operating Parameters
The variation in erosion rate (mg/min) of the 13Cr4Ni (13/
4) steel with time is shown in Fig. 4. From this figure, it
could be observed that for most of the cases, erosion rates
attained a steady state after 4th or 5th minute of testing.
Thereafter, irrespective of the case, almost a constant
erosion-rate-versus-time behaviour can be observed from
Fig. 4. The effect of velocity, impingement angle and sand
concentration on erosion rates is shown in Fig. 5. This plot
indicates that with the increase in velocity, the erosion rate
of 13/4 steel increases. The effect of velocity on the erosion
rates was calculated according to Eq. (1).
E ¼ KVn
ð1Þ
Here, ‘‘E’’ is the erosion rate in mg/min, ‘‘K’’ is the
proportionality constant, ‘V’ represents the velocity and
‘‘n’’ is the velocity exponent. The value of velocity
exponent (n) calculated according to Eq. (1) was found
to be around 0.95 for 13/4 steel. This value is very much
close to the one reported by Shivamurthy et al. [11].
Shivamurthy et al. [11] reported that the value of velocity
exponent for 13/4 steel is around 0.85. However, in
general, value of velocity exponent greater than 2 is
commonly reported in the literature. The low values of
exponent in both the cases could be related to the critical
energy required for the removal of material. As also
pointed out by Clark et al. [22], there is a critical value of
energy that is required for the removal of unit quantity of
material. Erodents possessing energy lower than critical
value might not be able to remove the material effectively.
As a result, the rise in velocity (below the critical value)
would not be much effective. This could be one possible
reason for the low value of exponent.
The 13/4 steel showed higher erosion rates at low impact
angle (30) as shown in Fig. 5. Similar results have also
been reported by other investigators [10, 15]. Concentra-
tion on the other hand has shown somewhat dubious trend.
Figure 5 indicates that at low velocity, an increase in the
concentration results in the decline in erosion rates. How-
ever, at high velocity, with an increase in the concentration,
the erosion rate increases. These results indicate high level
of interaction amongst the velocity and concentration as
also discussed in next section. It should be noticed that
with the rise in solid content, the probability of interaction
amongst the particles also increases.
The rebound velocity (Vout), the velocity with which the
particle would revert after impact, depends upon the
impinging velocity (Vin). It is known that restitution
coefficient (e) is inversely related to Vin [23–25]. This
indicates that as we increase the impact velocity, the
rebound velocity (Vout) decreases. Therefore, at low-
impact velocity, particles would be able to reflect more
efficiently than during high-velocity impacts. In the light of
above arguments, considering the case when the concen-
tration of particles is high enough to cause the interaction
between incoming and rebounding particles, the role
played by impact velocity cannot be ignored [26]. When
impacts take place at low impact velocity, the rebounding
particles would be able to de-accelerate and/or deviate the
incoming particles more effectively. However, when the
velocity of impact is high, the rebounding particles would
not be able to interact with the incoming particles with
Fig. 4 Variation in erosion rates with time of CA6NM steel
subjected to slurry erosion testing under different operating conditions
16
4
0.00
0.05
0.10
0.15
0.20
0.25
90
30
90
{
{
{
Erosion
rate
(mg/min)
Velocity (m/s)
Low concentration
High concentration
{
30
°
°
°
°
Fig. 5 Effect of different operating conditions on slurry erosion
performance of13Cr4Ni steel
292 Tribol Lett (2013) 52:287–303
123

8. similar intensity, as is the case during low-velocity
impacts. Papini et al. [27] proposed a computer simulation
model and carried out a detailed study to understand the
effect of nozzle geometry, impingement angle and velocity,
restitution coefficient and particle parameters on particle–
particle and particle–target interaction. It was reported that
restitution coefficients play an important role in deter-
mining the level of interference. They compared the results
of their model with the experimental results of Shipway
and Hutching [26] and found them to be in agreement. It
was also pointed out that at lower value of restitution
coefficient (high velocity) that rebounded particles do not
have enough velocity and tend to stay near the target sur-
face. This observation is in line with the concept forwarded
in present work as illustrated in the model shown in Fig. 6.
Papini et al. [27] further pointed out that shielding effect is
more prominent at high velocities (low restitution coeffi-
cient), an observation which differs from the observation in
present work. Authors here want to point out that in both
the studies (Papini et al. [27]; Shipway and Hutching [26]),
effect of fluid viscosity was not considered. These studies
were conducted with gas as the carrier fluid. However, in
present work, carrier fluid was a liquid (water). The drag
force on particle controlled by the viscosity of the fluid
would play a significant role in the erosion process. In case
of gas as a carrier medium, the rebounded particle, which
in case of high velocity could stay near to the target surface
(Fig. 6), could be carried away if liquid is used in place of
gas. This would reduce the effective shielding at high
velocities. Therefore, at low-velocity condition, shielding
effect provided by the rebounding particles would be more
profound in comparison with high-velocity condition when
carrier fluid is liquid. Hence, at low velocity, the erosion
rates would reduce with an increase in concentration.
However, at high velocity, erosion rates could increase
with concentration due to reduced shielding effect.
XRD analysis of the eroded and un-eroded samples is
shown in Fig. 7. The 13/4 steel was composed of mar-
tensite (a) and austenite (c) phase. The XRD analysis of the
tested samples under different erosive conditions indicates
the dissolution of c (310) peak. This trend was observed for
almost all samples tested under different testing conditions,
except for 30041 case. A significant variation in relative
intensities of the a phase peaks was also observed. The
relative intensities of peaks related to a (200) and a (220)
were reduced in comparison with un-eroded steel. These
results indicate that erosion process has affected the texture
of the target steel. The extent to which the texture was
affected would obviously depend upon the severity of the
erosion process. A validation of this argument that erosion
process has affected the texture of the target steel needed to
be undertaken using more sophisticated characterization
techniques such as electron back scatter diffraction
(EBSD). Overall, during erosion process, c phase being
soft in comparison with a phase might have got removed as
a result of impacts of erodent particles. Moreover, the
texture of the target steel was also affected to some extent.
3.1.2 Interaction Study
3.1.2.1 Line Plots Line plots showing the interaction
between different sets of parameters are shown in Fig. 8. The
Target surface
Incoming partilces
Rebounding partilce
Low velocity impact High velocity impact
elow
ehigh
e
c
a
f
r
u
s
t
e
g
r
a
T
e
c
a
f
r
u
s
t
e
g
r
a
T
Particles moving with
higher proportion of
initial kinetic energy
Interaction between
rebounded and incoming
particle
Due to low rebound velocity
particles are dragged away
by the fluid paving way for
fresh incoming particles
Fig. 6 Schematic diagram showing the effect of velocity on the shielding affect provided by the rebounding particles
Tribol Lett (2013) 52:287–303 293
123

9. level of interaction amongst the lines indicates the degree of
interaction amongst the variables. It is to be noticed that lines
in Fig. 8a are almost parallel to each other. This indicates
that the interaction between impingement angle and con-
centration (A 9 C) is almost negligible, whereas the plots
given in Fig. 8b, c indicate the interaction between
impingement angle and velocity, (A 9 V) and between
concentration and velocity, (C 9 V). The degree of inter-
action between V and C appears to be more evident than that
between V and A. For V 9 A interaction, it can be observed
that effect of increasing velocity is more prominent at low
impingement angle than at normal impingement angle. This
interaction is easily understandable as the components of the
velocity are directly related to the angle of impact. However,
the degree of interaction between V 9 A is thought to be
effected by the mechanism responsible for erosion. More
detailed discussion regarding the interaction effect between
V 9 A will be undertaken in section on erosion mechanism.
In case of V 9 C interaction, as discussed in the preceding
section, both velocity and concentration influence each other
significantly. From the slopes of the constant concentration
lines, it appears that at 0.5 wt% concentration, change in
velocity has more influence on erosion rate.
3.1.2.2 Contour Plots Contour plots showing the inter-
action between the variables are shown in Fig. 9. These
plots other than helping in identifying the interaction also
help in understanding the influence of individual parameter
on the interaction. Fig. 9a shows the interaction between the
concentration and impingement angle. It can be observed
that contour lines are almost parallel to each other, indi-
cating no interaction between the variables. Moreover, the
inclination of contour lines indicates that the variation in
erosion rates appears to be highly affected by the impinge-
ment angle in comparison with concentration. It also indi-
cates that maximum erosion is taking place at lowest impact
angle and highest concentration. In Fig. 9b, contour plot
showing the interaction between the impingement angle and
velocity is presented. It could be observed that with an
increase in impact angle, the contour lines are diverging
slightly. This indicates that interaction is becoming promi-
nent at higher impact angles. The large number of contour
lines along the velocity axis indicates that velocity is
affecting the erosion rates more significantly than impact
angle. The maximum erosion is observed for 30 angle and
maximum velocity condition.
The interaction between velocity and concentration is
given in Fig. 9c. The maximum level of interaction is
observed in this plot, as indicated by the slopes of the
contour lines. The interaction appears to be significant at
high concentrations in comparison with low concentra-
tions. It is further observed that both concentration and
velocity are playing significant role in interaction. In high-
velocity regime, concentration is significantly affecting the
erosion rates, whereas velocity is playing a more signifi-
cant role in low-velocity regime. Maximum erosion rates
are observed at highest values of concentration and
velocity.
Fig. 7 XRD profile of 13Cr4Ni
steel under un-eroded and
eroded conditions
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10. 3.2 Erosion Mechanism
3.2.1 Effect of Velocity
Macrographs showing the samples eroded at 30 and 90
impingement angles are shown in Fig. 10. It can be observed
that for 90 impingement angle, almost circular erosion scar
was formed; however, for 30 impact angle case, nearly
elliptical scar was produced. The SEM micrographs of the
eroded surfaces of 13/4 steel at two different velocities are
shown in Fig. 11. Other testing conditions for both the cases
were kept identical. The prominent erosion mechanism for
both the testing velocities was observed to be the formation
and removal of material in the form of platelets as shown by
high-magnification micrograph in Fig. 12. However, the
level of severity was obviously enhanced with velocity. In
case of low velocity, large number of platelets attached to
the eroded surface could be observed. However, when
velocity was increased, the number of platelets attached was
significantly reduced. The pit density calculated as number
of pits produced per unit area due to impact of particle was
calculated using ImageJ software. It was found that in case
of 4 m/s velocity, pit density was around 0.007/lm2
. How-
ever, when velocity was increased to 16 m/s, approximately
threefold rise in the pit density was observed. In latter case,
pit density was around 0.02/lm2
. This leads to the fact that
an increase in pit density resulted in a proportional rise in the
erosion rates. It was also observed that the surface eroded at
low-impact velocity was less plastically deformed in com-
parison with one eroded at high velocity. The removal of
material through platelet mechanism is comparatively a
slow process [28, 29]. It requires higher number of impacts
in comparison with microcutting or ploughing for the final
detachment of the material. After the formation of platelets,
the impacts by subsequent particle would result in flattening
of the platelets [30, 31]. Upon extreme flattening, the cracks
will be generating when strain exceeds critical value. This
process will eventually result in the removal of material in
the form of small fragments, indicating the presence of
fatigue phenomena. Along with platelet mechanism,
extensive plastic deformation was also playing a significant
role in the erosion process at high velocity. The intensive
Fig. 8 Line plots showing interaction between (a) angle of impingement and concentration (b) angle of impingement and velocity
(c) concentration and velocity, under slurry erosion condition for CA6NM steel
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11. indentation by the erodent particles at high velocity would
cause significant amount of plastic deformation of the sur-
rounding region as shown in Fig. 13. Due to this indentation
process, hardness of the material adjacent to the impact zone
is likely to increase due to work-hardening effect. The
subsurface hardness data of the eroded samples are shown in
Fig. 14. It could be observed that samples eroded at 16 m/s
velocity showed an increase in hardness irrespective of the
impingement angle. Samples eroded at 4 m/s velocity did
not show any observable increment in the subsurface hard-
ness. This might be due to the fact that the thickness of the
modified subsurface layer at 4 m/s velocity could be small
enough to be detectable at load and instrument utilized for
measuring the microhardness in the present case. At 16 m/s
Fig. 9 Contour plots showing interaction between (a) angle of impingement and concentration, (b) velocity and angle of impingement and
(c) velocity and concentration, under slurry erosion condition for CA6NM steel
Fig. 10 Macrographs of the 13Cr4Ni steel samples eroded at (a) 30 and (b) 90 impingement angles
296 Tribol Lett (2013) 52:287–303
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12. velocity, subsurface layer of about 30–40 lm beneath the
eroded surface appears to have been work-hardened due to
intensive plastic deformation caused by the impact of
erodent particles, which are now moving at higher velocity
in comparison with the particles travelling at around 4 m/s
velocity. At 16 m/s, velocity hardness of the subsurface
layer below the eroded surface showed an increase of around
15–20 % in comparison with un-eroded steel. An increase in
hardness below the eroded surface due to work-hardening
also helps in explaining the erosion mechanism (plastic
deformation) proposed in Fig. 15. With further continuation
of the impact process, brittleness of the material could
increase due to extensive work-hardening. This would result
in the removal of material from the surface in the form of
small fragments as illustrated in Figs. 13, 15.
3.2.2 Effect of Impingement Angle
The difference in erosion mechanism with change in
impingement angle is shown in Fig. 16. The surface
impacted at normal impingement angle showed the pre-
sence of platelets and plastically deformed target surface.
The impact energy of the particles would make the material
to deform plastically. This plastically deformed material
(a)
(b)
Platelets
Platelets
Platelets
Fig. 11 Morphology of slurry eroded surfaces of 13Cr4Ni steel at
(a) 4 m/s and (b) 16 m/s after 1 min of testing
Impact sites
Platelets
Fig. 12 High-magnification micrograph of the eroded 13Cr4Ni steel
showing the platelet mechanism of erosion
Fig. 13 SEM micrographs of 13Cr4Ni steel eroded at 16 m/s
velocity and 90 impingement angle after 2 min of testing
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13. would tend to flow outward and get accumulated around
the impact crater as shown in Figs. 12, 16. During sub-
sequent impacts, this deformed material would get
removed in the form of small fragments as discussed in
preceding subsection. For surfaces impacted at acute angle,
major material removal mechanisms were ploughing and
mixed cutting–ploughing mode. Hutching [32] and Levy
[31, 33, 34] have shown that with the impact of round
particles such as sphere, the material displaces and get
accumulated at the end from where the particle leaves. This
mechanism of material removal is generally known as
ploughing. However, in our work, in addition to ploughing
observed at acute impingement angle, the primary mode of
material removal was the mixture of ploughing and mi-
crocutting mechanism. This mechanism is significantly
different from what the earlier researchers have proposed,
which the authors have named to be mixed cutting and
ploughing (Mcp). Authors believe that Mcp is the effect of
the particle shape as illustrated in Fig. 17. Pure microcut-
ting was observed to be secondary in nature. Ideal micro-
cutting and ploughing are shown in Fig. 17 with the help of
illustration. It should be taken into consideration that in
case of ideal cutting, material will be removed from the
surface through shearing action. The volume of material
would be sheared out by very sharp particle as shown in
Fig. 17a. The chances that debris of material will remain
attached with the crater are minimal. However, for pure
10 20 30 40 50 60
350
375
400
425
450
475
500
Sub-surface
hardness
of
eroded
samples
(HV
0.01Kgf
)
Distance from top surface (µm)
2045 9045
20165 90165
Base hardness :390 HV0.01 Kgf
Fig. 14 Subsurface hardness of the slurry eroded 13Cr4Ni steel
samples
Indent
Erodent particle
Target surface
Crack formed after extreme work-hardening
Further initiation of cracks
with loading and work-hardening
Detached Fragment
Highly plastically deformed region
Fig. 15 Schematic diagram illustrating the plastic indentation mech-
anism of erosion
(a)
(b)
Fig. 16 SEM micrographs showing the effect of impingement angle
on the erosion mechanism of 13Cr4Ni steel eroded at 16 m/s,
0.1 wt% sand concentration and (a) 30 (b) 90 impingement angle
298 Tribol Lett (2013) 52:287–303
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14. ploughing case, displacement of the material could be
imagined to have taken place by a blunt object such as
sphere. The material in this case would be plastically
deformed and displaced. Most of the displaced material
would accumulate at the exit point of the particle.
When particle shape is in-between to that of extremely
sharp and blunt type, the erosion process is likely to be
taken place involving both the microcutting as well as
ploughing. Figure 17c represents a case of mixed cutting–
ploughing mechanism. In this case, the removal of material
would be initiated through cutting process; however, due to
the subangular shape of the particles as used in the present
study, the deformation of material would proceed plasti-
cally. Therefore, along with initial shearing, plastic defor-
mation is likely to take place. This represents a mixed
cutting–ploughing (Mcp) mode. The SEM micrographs
Fig. 17 Schematic diagram
illustrating the effect of erodent
shapes on erosion mechanism of
ductile material at low
impingement angles, (a) ideal
microcutting, (b) ideal
ploughing and (c) mixed
cutting–ploughing mode
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15. shown in Fig. 16 clearly indicate the presence of Mcp
mode. Moreover, the presence of this mode was more
preferential than pure ploughing or microcutting.
In continuation to the discussion of Sect. 3.2.1, the
reason for the lower erosion rates for normal impingement
angles in comparison with acute angle impingement could
be easily explained in terms of erosion mechanism. Platelet
mechanism as explained earlier is a slow process involving
combined plastic deformation and fatigue phenomena. In
contrast to it, microcutting and Mcp are more efficient
mechanisms for the removal of material. This explains the
cause as to why the erosion rates were higher at low
impingement angles rather than at normal impact angle for
13/4 steel.
3.2.3 Effect of Slurry Concentration
The effect of concentration on the erosion mechanism can
be observed from Fig. 18. The comparison of micrographs
indicates no significant difference of concentration on
erosion mechanism. The material removal took place
through platelet mechanism at normal impacts and
ploughing along with Mcp at acute impingement angle.
However, it is to be noticed that intensity of the impacts
was significantly affected. At low velocity, large number of
impacts could be observed for low concentration; however,
with an increase in concentration, the intensity seems to be
reducing. For instance, the average pit density for 4 m/s
velocity at 0.1 and 0.5 wt% sand concentration was 0.007
and 0.004/lm2
, respectively. These values are the average
of five consecutive values for different locations. The
variability in the pit density was of the order of ±0.0008/
lm2
. At high velocity, the trend was opposite to that
observed for low velocity. This observation appears to be
supporting the trend in erosion rates as discussed in Sect.
3.1.1.
3.2.4 Effect of Time
The erosion mechanism was not influenced by time during
the time interval investigated in this present study. Similar
erosion mechanism was observed after 1 and 10 min of
testing. In case of normal impingement angle, formation
and removal of platelets through subsequent impacts was
observed to be the primary erosion mechanism even after
10 min of testing. However, the eroded surface was
observed to be in highly plastically deformed state at the
end of experimentation. In case of low-angle impingements
also, erosion mechanism essentially remained same even
after 10 min of testing. This observation comes in support
of Fig. 4, wherein constant erosion rates were shown from
initial to final stage of experimentation.
3.2.5 Effect of Distance
The variation in erosion mechanism with radial distance
from the impact zone was readily observed. For low impact
angle, the signs of normal impacts at the centre of the
impact zone were clearly visible as shown in Fig. 19a.
Moving away from the impact zone resulted in the change
of mechanism. Now, the craters were elongated with the
presence of ploughing and Mcp as the operative mecha-
nisms. Moving further away from the impact zone resulted
in an reduction in the intensity of the impacts (Fig. 19b);
here, ploughing along with microcutting was the mecha-
nism to be observed.
At normal impingement angles, the core of the impact
zone showed the presence of large number of normal
impact. The platelet mechanism was observed in this zone.
Moving radially outward from the core of the impact zone,
presence of ploughing and Mcp was observed after a par-
ticular distance as shown in Fig. 19c. This distance was
found to be the function of impact velocity. A particle
(a)
(b)
Fig. 18 SEM micrographs showing the effect of concentration on
erosion mechanism of 13Cr4Ni steel eroded at 4 m/s, 90 impinge-
ment angle and (a) 0.1 wt% (b) 0.5 wt% sand concentration
300 Tribol Lett (2013) 52:287–303
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16. moving in a fluid would be acted upon by two different
force, viz. inertia force and drag force. Drag force of the
fluid near the impact zone would tend to deviate the
particle from its path, whereas inertial force would resist it.
With the rise in velocity, increase in drag force would be
more profound than an increase in inertia of the particle. As
a result of this, particles in a jet of low velocity would have
lesser tendency to deviate than those in a jet of high
velocity. Thus, the radius of the core of the jet, wherein
particles are travelling normally to the target surface,
would reduce with an increase in velocity.
3.3 Statistical Analysis
Using the experimental results obtained from this study, a
mathematical model was developed using regression
approach. In the developed model shown in Eq (2), V, C,
and Ai represent velocity in m/s, concentration in wt% and
angle of impingement in degrees, respectively. K1 to K6 are
constants, the values of which are given in Table 4. The
developed model shown in Eq (2) was able to predict the
erosion rates with an error of ±2 %.
Erosion rate ¼ K1 þ K2V K3Ai K4C K5VAi
þ K6CV ð2Þ
Table 4 Values of constants of
regression model used for pre-
dicting slurry erosion rates of
CA6NM steel
Constant Value
K1 0.152
K2 6.35 e-03
K3 3.32 e-04
K4 9.04 e-02
K5 2.22 e-05
K6 9.79 e-03
(a)
(b)
(c)
Fig. 19 SEM micrographs showing the variation in erosion mecha-
nism of 13Cr4Ni steel with radial distance from the centre of impact
zone for the case (a) 30 impingement angle, 4 m/s velocity at the centre
of impact zone, (b) 30 impingement angle, 4 m/s velocity at a distance
of 8 mm from the centre of impact zone and (c) 90 impingement angle,
16 m/s velocity at a distance of 1.2 mm from the impact zone
Fig. 20 Comparison of the predicted and actual erosion rates for
CA6NM steel from the presented study
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17. The comparison between predicted and actual erosion
rates is shown in Fig. 20. A linear trend between the
predicted and actual erosion rates could be observed. The
validity of the model was checked using the experimental
results from the published literature. Figure 21 shows the
comparison between the predicted and actual results
obtained from the literature. It could be seen that within
acceptable error range of 12 %, model was able to predict
the erosion rates of Manisekaran et al. [10]. However, the
error in case of results given by Sugiyama et al. [13] was
greater than 50 %, whereas model was able to predict the
results of Shivamurthy et al. [11, 12] with an error of 30 %.
Moreover, the model was not able to predict accurately the
results presented by Sugiyama [13]. This discrepancy
between the predicted and actual results in case of
Sugiyama et al. [13] could be related to the size of
particles utilized by the authors. Sugiyama et al. [13]
employed a particle size of around 80 lm for their studies.
Therefore, for further improvement in this model,
experimental results from literature could be utilized for
the development of regression model, which would help in
enhancing the predictability of the model.
4 Conclusion
Effect of various operating parameters on slurry erosion
mechanism of CA6NM steel was investigated. The fol-
lowing are some of the important conclusions drawn from
the investigation,
• The erosion mechanism was significantly affected by
the velocity and angle of impingement. At 90
impingement angle, platelet mechanism was observed
to be the dominant erosion mechanism, whereas at 30
angle, ploughing along with proposed mixed cutting–
ploughing mechanism was observed to be responsible
for the removal of the material.
• Erosion mechanism was not affected by other param-
eters such as concentration and testing time; however,
the pit formation intensity was surely affected.
• Erosion mechanism changes with the radial distance
from the impact zone. For normal impact angle, the
removal of material at the centre of the impact zone
took place through platelet mechanism; however, at the
periphery of the impact zone, ploughing and microcut-
ting were responsible for erosion of the material.
• At high velocity, other than platelet mechanism,
removal of material in the form of fragments through
plastic indentation mechanism was also observed. A
model has also been proposed for the observed
mechanism.
• Interaction between velocity and concentration was
most significant.
• Austenitic phase was more prone to erosion in
comparison with martensite phase as confirmed by
XRD analysis.
• A regression-based model developed using experimen-
tal results was able to predict the erosion rates of the
present study as well as from literature, appreciably
with a maximum error of 30 %.
Acknowledgments Authors thankfully acknowledge the financial
assistance provided by Council of Scientific and Industrial Research
(CSIR), India, under project title ‘‘Development of Slurry Erosion
Resistant Coatings for Hydroturbines’’, File no.: 22(0604)/12/EMR-II.
References
1. Humphrey, J.A.C.: Fundamentals of fluid motion in erosion by
solid particle impact. Int. J. Heat and Fluid Flow 11(3), 170–195
(1990)
2. Mann, B.S., Arya, V.: Abrasive and erosive wear characteristics
of plasma nitriding and HVOF coatings: their application in
hydro turbines. Wear 249(5–6), 354–360 (2001)
3. Bajracharya, T.R., Acharya, B., Joshi, C.B., Saini, R.P., Dahlh-
aug, O.G.: Sand erosion of Pelton turbine nozzles and buckets: A
case study of chilime hydropower plant. Wear 264(3–4), 177–184
(2008)
4. Iwabuchi, Y., Sawada, S.: Metallurgical characteristics of a large
hydraulic runner casting of type 13Cr-Ni stainless Steel. In: 1982,
pp. 332–354. ASTM
5. Antunes, F.F., Cornman, R.E., Hartkopf, R.J.: Centrifugal pumps
for desalination. Desalination 38, 109–122 (1981)
6. Jeske, H.O.: Charge-gas compressors in coal gasification and
olefin plants—25 years of experience with a vital plant compo-
nent. Chem. Eng. Process. 18(2), 113–122 (1984)
7. Kim, J.W., Kim, Y.S., Park, C.Y.: Failure analysis of cracking at
volute tongues of feedwater pump casings. Eng. Fail. Anal. 9(1),
17–30 (2002)
8. Chauhan, A.K., Goel, D.B., Prakash, S.: Erosion behaviour of
hydro turbine steels. Bull. Mater. Sci. 31(2), 115–120 (2008)
Fig. 21 Comparison of the predicted and actual erosion rates for
CA6NM steel from literature
302 Tribol Lett (2013) 52:287–303
123

18. 9. Grewal, H.S., Bhandari, S., Singh, H.: Parametric study of slurry-
erosion of hydroturbine steels with and without detonation gun
spray coatings using taguchi technique. Metall. Mater. Trans. A
43(9), 3387–3401 (2012)
10. Manisekaran, T., Kamaraj, M., Sharrif, S.M., Joshi, S.V.: Slurry
erosion studies on surface modified 13Cr-4Ni steels: Effect of
angle of impingement and particle size. J. Mater. Eng. Perform.
16(5), 567–572 (2007)
11. Shivamurthy, R.C., Kamaraj, M., Nagarajan, R., Shariff, S.M.,
Padmanabham, G.: Influence of microstructure on slurry erosive
wear characteristics of laser surface alloyed 13Cr–4Ni steel.
Wear 267(1–4), 204–212 (2009)
12. Shivamurthy, R.C., Kamaraj, M., Nagarajan, R., Shariff, S.M.,
Padmanabham, G.: Slurry erosion characteristics and erosive wear
mechanisms of Co-based and Ni-based coatings formed by laser
surface alloying. Metall. Mater. Trans. A 41(2), 470–486 (2009)
13. Sugiyama, K., Nakahama, S., Hattori, S., Nakano, K.: Slurry wear
and cavitation erosion of thermal-sprayed cermets. Wear
258(5–6), 768–775 (2005)
14. Romo, S.A., Santa, J.F., Giraldo, J.E., Toro, A.: Cavitation and
high-velocity slurry erosion resistance of welded Stellite 6 alloy.
Tribol. Int. 47, 16–24 (2012)
15. Sugiyama, K., Harada, K., Hattori, S.: Influence of impact angle
of solid particles on erosion by slurry jet. Wear 265(5–6),
713–720 (2008)
16. Sugiyama, K., Harada, K., Hattori, S.: Prediction of the volume
loss by using slurry jet test on SCS6. J. Solid Mech. Mater. Eng.
2(7), 955–966 (2008)
17. Mann, B.S.: High-energy particle impact wear resistance of hard
coatings and their application in hydroturbines. Wear 237,
140–146 (2000)
18. Chattopadhyay, R.: High silt wear of hydroturbine runners. Wear
162, 1040–1044 (1993)
19. Santa, J., Baena, J., Toro, A.: Slurry erosion of thermal spray
coatings and stainless steels for hydraulic machinery. Wear
263(1–6), 258–264 (2007)
20. Santa, J.F., Espitia, L.A., Blanco, J.A., Romo, S.A., Toro, A.:
Slurry and cavitation erosion resistance of thermal spray coatings.
Wear 267(1–4), 160–167 (2009)
21. Grewal, H.S., Agrawal, A., Singh, H.: Design and development of
high-velocity slurry erosion test rig using CFD. J. Mater. Eng.
Perform. 22(1), 152–161 (2013)
22. Clark, H.M., Hawthorne, H.M., Xie, Y.: Wear rates and specific
energies of some ceramic, cermet and metallic coatings deter-
mined in the Coriolis erosion tester. Wear 233–235, 319–327
(1999)
23. Stack, M.M., Corlett, N., Zhou, S.: Some thoughts on the effect of
elastic rebounds on the boundaries of the aqueous erosion-cor-
rosion map. Wear 214(2), 175–185 (1998)
24. Tabor, D.: The hardness of metals. Oxford University Press, USA
(2000)
25. Hertz, H.: On contact of elastic solid. In: Miscellaneous papers,
vol. 3. London: Macmillan, New York, Macmillan and co.,
(1896)
26. Shipway, P.H., Hutchings, I.M.: A method for optimizing the
particle flux in erosion testing with a gas-blast apparatus. Wear
174(1–2), 169–175 (1994)
27. Papini, M., Ciampini, D., Krajac, T., Spelt, J.K.: Computer
modelling of interference effects in erosion testing: effect of
plume shape. Wear 255(1–6), 85–97 (2003)
28. Hutchings, I.M.: A model for the erosion of metals by spherical
particles at normal incidence. Wear 70(3), 269–281 (1981)
29. Rickerby, D.G., MacMillan, N.H.: The erosion of aluminum by
solid particle impingement at normal incidence. Wear 60(2),
369–382 (1980)
30. Levy, A., Aghazadeh, M., Hickey, G.: The effect of test variables
on the platelet mechanism of erosion. Wear 108(1), 23–41 (1986)
31. Levy, A.V.: The platelet mechanism of erosion of ductile metals.
Wear 108(1), 1–21 (1986)
32. Hutchings, I.M., Winter, R.E.: Particle erosion of ductile metals:
A mechanism of material removal. Wear 27(1), 121–128 (1974)
33. Bellman Jr, R., Levy, A.: Erosion mechanism in ductile metals.
Wear 70(1), 1–27 (1981)
34. Levy, A.V., Chik, P.: The effects of erodent composition and
shape on the erosion of steel. Wear 89(2), 151–162 (1983)
Tribol Lett (2013) 52:287–303 303
123
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