45

3. Material wear characteristics

3.1 Introduction

Erosion occurs when particles in the fluid stream impact the pump hydraulic surfaces and

cause material removal. Erosion wear is the major type of wear in slurry pumps.

Understanding erosion involves two levels of investigation: a) the determination of particle

trajectories in the fluid stream and b) the mechanisms by which the particles remove

materials. This section examines both of these areas as they pertain to slurry pumps and

typical slurry pump materials.

From a micro-mechanical viewpoint, the wear process (whether it is erosion or abrasion)

takes place due to a wide range of different wear mechanisms. Some of these are outlined

by Zum Gahr (1987) and Roco (1990) and may include: adhesion, micro-machining,

micro-cutting, fatigue, fretting, delamination, plastic deformation and rolling wear. Material

removal mostly happens due to a number of separate wear mechanisms occurring

simultaneously. The specific impact conditions, as well as the material and particle

characteristics, determine the extent of material removal.

The significance of a particular wear mechanism in a slurry pumping situation depends on

the type of material used for the wetted parts. The different classes of materials (metals,

elastomers, glasses, ceramics and composites) all have different erosion wear

characteristics. Specific slurry conditions can produce very different wear rates in the

different material classes. Selecting the best material option is a key task in maximising the

wear life of slurry pumps.

This section reviews the published literature with respect to the basic physical properties

and characteristics of cast white iron and elastomer (including natural rubber and

polyurethane). Laboratory wear tests results are also compared.

46

A detailed review of wear test methods has been undertaken to choose the best test.

Justification is provided for using the slurry jet tester to determine wear characteristics.

Finally, the wear results from a wide range of tests carried out on a number of typical white

irons and elastomers are presented and analysed.

3.2 Erosion wear process

3.2.1 Defining erosion

Wear is basically a material removal process that causes a (generally) gradual deterioration

or removal of (generally) the surface on which the process is acting. The ASTM (1983)

definition of erosion wear covers material removal from a solid surface due to mechanical

interaction between the surface and impinging particles in a liquid stream. This differs

from abrasion where the particles are forced against and move along the solid surface.

The essential difference between the two is that erosion involves the transfer of kinetic

energy to the surface and abrasion does not. Hutchings (1992) takes the definition one

step further making the distinction between slurry erosion (particles in a liquid) and the

more common erosion occurring due to particles in a gaseous stream.

Roco (1990) states that erosion wear is the principal concern in the operation of the slurry

pump after energy consumption. Wilson et al. (1992) explain that erosive wear of the

wetted parts of the pump limits the useful life. Impingement erosion wear occurs at the

leading edge of impeller vanes, the casing-liner cutwater and at significant changes in flow

direction (Roco, 1990). Abrasion is not so significant and only occurs at the shaft sleeve

and occasionally between the impeller and the side-liners involving large particle slurries.

Other factors that can have a profound effect on the magnitude of overall erosion include;

corrosion and cavitation, which when present, can act synergistically to increase wear rates

(Roco, 1990).

47

3.2.2 Types of erosion wear

In slurry erosion, energy is transferred from the particle to the material due to the force of

the particles impact velocity and momentum. In many cases this is sufficient to cause

material removal.

Roco (1990) outlined several forces of different origins that may act on the particle. These

include:

- impact from neighbouring particles (contact forces)

- drag due to the flowing fluid (drag forces), and

- gravity (gravity forces).

In a curving and rotating flow field (such as the pump impeller vanes, casing-liner periphery

or impeller back shroud) Wilson et al. (1992) described two additional forces due to:

- centrifugal acceleration (centrifugal force), and

- Coriolis acceleration (Coriolis force).

The erosion wear rate depends on the number, mass and shape of individual particles

striking the surface and their impact velocity. Particles entrained in a slurry may impact the

surface at any angle from 00 to 900. The angle of impingement largely determines the type

of wear.

Roco (1990) identified three major wear modes. These are illustrated in Figure 3.1 and

include:

a) directional impingement,

b) random impingement, and

c) sliding bed friction.

48

Wilson et al. (1992) likened the sliding bed friction to the contact bed portion of flow in a

pipe. The contact load portion remaining unsuspended by turbulence.

Figure 3.1 Three different modes of slurry erosion wear a) direct impingement, b)random

impingement, c) sliding bed friction (after Roco, 1990).

Near the material surface in dense slurry flow, Roco (1990) argues that the particle-particle

interaction stresses are at least of the same order of magnitude as the liquid and liquid-

particle stresses. The particles are then in either turbulent suspension or supported by

contact with other particles and will interact with the material when the convective velocities

are directed towards the material, or the particle is one of a group of particles sliding along

the material.

49

Roco used a modified Froude (Fr*) number, which is the ratio of turbulent dispersive force

to inertial force acting on the particle, to determine the predominant wear type in a pump.

A Froude number less than or equal to one means that the inertial forces are greatest and

that the frictional particle-particle interaction plays the most important role.

In flows where the radius of curvature is small, the centrifugal acceleration can be very high.

Centrifugal acceleration is equal to V2/r, where V is the tangential velocity and r is the

radius of curvature (Wilson et al., 1992). This relationship has serious implications for both

sliding bed and impingement wear with changes in flow direction. It is also very important

for vortical flows and in situations where strong turbulent eddies exist. As stated previously

in Section 2, Pogodaev and Tsvetkov (1994) put forward that the local wear rate is

proportional to the vortex velocity cubed.

3.2.3 Stochastic nature of erosion

Shook and Roco (1991) state that the erosion process will vary as a function of time and

location. Particles of generally random size and shape slide, roll and impact onto a

microscopically non-uniform material surface with random vector velocity under various

impingement angles. The amount of material removed (Ds) from the surface can only be

determined statistically. A typical wear histogram is shown in Figure 3.2.

50

Figure 3.2 Typical wear histogram (Shook and Roco, 1991)

For similar reasons, the energy of the impinging particle (MV2/2) can also only be

determined statistically. Roco (1990) states the energy has a probability distribution

function characterised by a variance which is about the same order of magnitude as the

mean value for flows in a centrifugal pump. The wear rate increases with the variance of

the probability distribution for any given mean, due to the presence of a greater number of

larger instantaneous values.

3.2.4 Erosion wear model

Mutton (1988) has outlined a range of factors which make up the abrasion wear system

(in this particular text abrasion was defined to include erosion). They are grouped into

three categories including:

- the properties of the wear material,

- the properties of the abrasive material,

- the nature and severity of the interaction between abrasive and wear material.

Muttons wear variables are listed in Table 3.1. The material properties have been

separated out into the two material groups (white irons and elastomers).

51

Table 3.1 Factors which influence erosive wear behaviour (modified from Mutton, 1988)

Contact

Conditions

Particle

Properties

Material

Properties

Velocity

Impingement angle

Concentration

Temperature

Shape

Size

Hardness

Friability

Yield strength

Fracture

White iron:

- hardness

- elastic modulus

- ductility

- fracture toughness

- microstructure

Elastomer:

- tear resistance

- fatigue strength

- tensile strength

- elastic modulus

- resilience

- hardness

- elongation

To further aid in understanding the erosion process a model is proposed by the author

based on the above work of Mutton, but also including the surrounding fluid properties.

The fluid carrier properties (pH, chemical composition, temperature, volume of dissolved

gases, rheology) ultimately determine the contact conditions and even the material

properties to some extent. The proposed model for erosion wear is shown in Figure 3.3.

52

Figure 3.3 Erosion wear model

3.3 Pump wear materials

Cast white irons were the first material to be used to combat the wear experienced in the

originally soft grey iron centrifugal slurry pumps. After the development and introduction of

cast white irons in the early 1900s, the new material became widely used. The next major

Fluid characteristics (rheology, density,velocity)

Particle characteristics

(size, shape, mass, hardness)

Contact conditions (velocity, angle, no. of impacts)

Material properties (hardness, tear strength, fracture,

toughness, etc)

53

development in the 1930s, was the use of gum rubber for slurry pump liners. White iron

and rubber remain the materials of choice for most slurry pump manufacturers today.

3.3.1 White Iron

Huggett (1992) described white iron as an alloy of iron containing no free graphitic carbon

and which when fractured has a white fine crystalline fracture occurring along the carbide

plates. They are characteristically hard (500 HBN plus), not very tough and may fracture

catastrophically if subjected to high stress. The eutectic and hypereutectic types of white

irons in the as-cast state are multiphase alloys that contain hard carbides within a typically

austenitic ferrous matrix. Further, they are typically high alloy irons with high levels of

chromium and possibly nickel, molybdenum and/or manganese. The ferrous matrix of the

alloy is usually austenite, with martensite, pearlite and bainite possible depending on thermal

history.

The erosion wear resistance of a white iron has been found to depend on the

microstructure and physical characteristics (Walker and Huggett, 1990).

3.3.1.1 White iron microstructure

For an iron-chromium-carbon (Fe-Cr-C) alloy system the four primary phases present on

solidification are austenite, ferrite, (Fe,Cr)7C3 carbide and FeC3 carbide (Thorpe and

Chico, 1985).

The relative amounts of Fe, Cr and C determines the phase structure of the white iron.

Two main structures are of interest: Eutectic and Hypereutectic. These are usefully

summarised by Huggett (1992).

a) Eutectic white irons (Material code A05)

54

In this type of alloy the iron remains in its molten state until the eutectic point is reached and

solidification of carbide and austenite phases occurs simultaneously. This results in an

interlaced structure of continuous carbide. If the composition lies directly on the eutectic

point then the carbide microstructure is extremely fine and difficult to resolve.

b) Hypereutectic white irons (Material code A12, A211, A217)

On solidification, this alloy consists of primary (Fe,Cr)7C3 carbides in a eutectic matrix.

The primary carbides continue to grow until arrested at the eutectic point when the matrix

solidifies. The primary carbides are shaped like hexagonal rods with a high aspect ratio

and are generally discontinuous and quite coarse. Morphology and size of the primary

carbides are controlled by the thermal history. The volume fraction of primary carbides is a

function of the carbon content (for any given Fe and Cr).

Heat treatment is almost universally carried out on white irons to enhance hardness and/or

toughness. The hardening treatment involves heating to around 1000 0C, holding while

homogenisation occurs and then cooling at a sufficient rate to ensure that the austenite is

fully transformed to martensite without the formation of pearlite.

The typical microstructures of the two different types of white iron are shown in Figure 3.4.

(a) Eutectic (Mat. code A05) (b) Hypereutectic (Mat. code A12)

Figure 3.4 Typical white iron alloy microstructure (Walker and Huggett, 1990) mag. 300x

55

3.3.3.2 White iron physical properties

A comparison of the physical characteristics of the different white irons used in the test

work and compared to grey iron, are shown in Table 3.2. An important aspect of white

irons is that they are heterogeneous. At a macro level the material has one set of properties

while at the micro level each of the various phases have a different set of properties.

Table 3.2 Physical properties of cast irons used in test work

(Warman Material Data Sheet, 1994)

(* inoculated development alloys, refer Dolman et.al., 1998)

Material code Type Carbide Vol.

(%)

Hardness

(HBN)

Composition

A05 27% Cr white

iron to AS 2027

30 650 2.8C, 27Cr,

0.5Mo, Fe bal.

A12 30% Cr

Hypereutectic

50 650 30Cr

A218/211/217* Hypereutectic 50+ 700+ 4-5.5C, 27-37Cr,

1-4Mn, Fe bal.

G47 Heat treated

cast iron

0 110 3.5C, 2.0Si,

Fe bal.

Significant improvements in hardness and toughness of Hypereutectic white irons has been

achieved by inoculating the molten iron with fine carbide particles to refine the carbide size

(Dolman et al., 1998). This represents one of the latest approaches in metallurgical

development of white iron properties.

3.3.2 Elastomer

56

Clemitson (1997) defines elastomer as a low Elastic Modulus polymer such as rubber or

soft polyurethane. Elastomer is predominantly used in slurry pumps in applications that

have fine particles. Natural soft rubber is the material of choice for fine erosive and

corrosive slurries. In applications with high tip speeds or in the presence of large particle or

tramp materials, polyurethane is preferred (Wilson et al., 1992).

The typical mechanical properties for rubber and polyurethane are outlined by Mutton

(1988) in Table 3.3.

Table 3.3 Comparative mechanical properties of different elastomers (after Mutton, 1988)

Type Hardness

(Duro A)

Elastic Modulus

(MPa)

Tensile

Strength

(MPa)

Elongation

At Break

(%)

Nat. Rubber 45 75 0.7 3.5 10 20 700 - 900

Polyurethane 60A 60D 14 20 20 60 300 - 500

3.3.2.1 Natural rubber properties

A common rubber used in slurry pumps is a soft natural rubber crepe with additions of

different fillers (including carbon black for strength), anti-oxidants, waxes and accelerators

(Warman Material Data Sheet, 1994). The natural rubber consists of isoprenes with the

cis-1,4 configuration. The polymer is cross-linked at the double bonds to create a three

dimensional network. The linked polymer chains allow the rubber to retain its shape after

extension. Adjusting the degree of cross-linking and also the filler level controls the

hardness of the rubber. Mutton (1988) states that the typical hardness for rubber

components in slurry applications is 45-55 Durometer A. The material properties of the

specific rubbers used in this work are given in Table 3.4.

Table 3.4 Physical properties of rubbers used in test work

(Warman Material Data Sheets, 1994)

57

Rubber

Material

Code

Hardness

(Duro A)

Resilience

(%)

300%

Modulus

(MPa)

Tensile

Strength

(MPa)

Elongation

At Break

(%)

R26 43 82 1.5 21.5 450

R33 44 77 3.2 23.5 610

R24 50 - 3.0 20.0 600

R08 55 65 6.5 22.5 450

3.3.2.2 Polyurethane properties

Polyurethane can be formulated to give a wide range of different properties. Clemitson

(1997) outlines the three main ingredients in castable polyurethanes used in slurry pumps

and their effect on properties:

- Polyols - the backbone or soft segment

- Di Isocyanates - the hard or rigid phase

- Chain extender - an aromatic diamine with a chloro or methyl group added to control

the rate of reaction.

The polyurethane used in the work here had a polyether soft segment, TDI hard phase and

a MOCA cure. The physical properties are shown in Table 3.5.

58

Table 3.5 Physical properties of polyurethane used in test work

(Warman Material Data Sheet, 1995)

Poly

Material

Code

Hardness

(Shore A)

Resilience

(%)

300%

Modulus

(MPa)

Tensile

Strength

(MPa)

Elongation

At Break

(%)

U01 80 65 10.3 34.0 490

3.4 White iron wear characteristics

3.3.1 White iron mechanisms of wear

There appears to be very little published specifically on the wear mechanisms for white

irons. In practice, because of their heterogeneous nature, they fit somewhere between

metals and ceramics. They have a carbide phase with physical properties similar to

ceramic. The matrix phase has physical properties the same as martensitic steel.

The wear mechanisms for metals and ceramics are outlined by Lancaster (1990) in terms of

a deformation mode and particle detachment process. These are shown listed in Table 3.6.

Table 3.6 Wear mechanisms for different materials (Lancaster, 1990).

Material Deformation Mode Particle Detachment Process

Metal Plastic-Elastic Plastic Grooving Prow formation Cutting (chip formation)

Ceramics Elastic Brittle fracture

Crack propagation Flaking Fatigue

59

Similar to Lancaster, Hutchings (1992) proposes two major mechanisms for erosive wear.

These are plastic deformation and brittle fracture. The plastic deformation mode is most

applicable to softer metals. The fracture mechanism is important for metals that are hard or

brittle.

The wear process for metals is described by Hutchings (1992). The hard particle impacts

a soft ductile material surface, plastic deformation generally occurs and an indentation is

produced. Wear material displaced may form a rim of deformed material around the

indent or it may be removed as wear debris. In a simple analysis of the erosion process a

linear relationship exists between the mass of material removed from the surface and the

mass of material that has struck the surface.

Finnie (1995) shows the deformation shape of the material surface after the impact of a

hard particle. The amount of material removed depends on the impact velocity, the shape

and orientation of the particle and the impact angle. Rounded particles deform the surface

by ploughing (ie. displacing) material in front of, and to the side, of the indent, with little or

none being directly removed. Wear only occurs when some of the displaced material from

the indent is fractured with subsequent impacts and is ultimately lost. When more angular

particles contact the material surface the deformation which occurs depends on the

orientation of the particle. Figure 3.5 illustrates how the different impingement angle

influences the wear scar appearance.

60

Figure 3.5 Typical impact site cross-sections for different impingement angle

(Finnie, 1995).

Wang et al. (1993) show how the wear mechanisms are different for the carbide and matrix

phases with a tungsten white iron. They propose a complex wear mechanism based

around a number of steps including:

a) scratching of the matrix by the erodent

b) grain boundary is weakened by electrochemical corrosion

c) matrix is removed by repeated impacts

d) vortices form downstream of the carbides with cavitation aggravating matrix wear

e) exposed carbides crack, break and fall off under repeated impact.

This wear process is illustrated in Figure 3.6.

61

Figure 3.6 Tungsten white iron wear process (Wang et al., 1993)

3.4.2 Effect of microstructure and physicals on wear

The rate of erosive wear of white irons is dependant on the overall microstructure and in

particular the inter-carbide spacing (Day, 1982). Huggett (1992) believed the primary

factors were the carbide size, the matrix hardness and the relative volume fraction of

carbide and matrix. Optimum wear resistance was observed for alloys having a maximum

62

matrix hardness and minimum matrix volume fraction. Decreasing the carbide size also

resulted in a reduced wear rate. Carbide size was particularly important with sharp particle

slurries.

Dolman et al. (1998) also showed the effects of carbide size on wear in a hypereutectic

white iron. Figure 3.7 shows a wear rate reduction of some 20% with a reduction in

carbide size from 80 to 30 mm.

Figure 3.7 Effect of carbide size on wear rate (Dolman et.al., 1998)

The reduction in carbide size also has an effect on overall hardness with an increase of 100

HBN points in the example quoted above.

3.4.3 Effect of particle properties on wear

3.4.3.1 Effect of particle hardness

63

Mutton (1988) states that the hardness of the particles relative to the hardness of the

material surface (He/Hm) is the key parameter determining abrasive wear. It is possible

that at the particle scale, similar mechanisms would apply for erosion wear as well.

The hardness of some typical minerals and ores are given in Table 3.7. As Mutton (1988)

points out, it is not only the hardness of the mineral but that of the gangue (or waste)

minerals that might also be mixed in the ore that determines the wear rate.

Table 3.7 Properties of various minerals and ores (after Mutton, 1988).

Ore Mineral/Type Relative

Density

Hardness

(HV)

Compressive

strength

(MPa)

Bond

Work

Index

Bauxite

Coal

Limestone

Heavy Sulphides (Lead/Zinc)

Copper ore (Chalcopyrite)

Haematite

Granite

Quartz

4.9-5.2

1.2-2.0

2.3

7.4-7.6

4.1-4.3

5.3

-

2.7

150-420

150-250

150-250

-

350-400

470-650

500-800

800

-

5-40

130-200

60-100

-

180-250

100-300

140-650

-

12

12

11

12

9

17

17

Hutchings (1992) on reviewing abrasion wear of ductile metals and ceramics found that if

the material surface is more than 1.2 times harder than the erodent then the wear rate is

substantially reduced. This effect is illustrated in Figure 3.8.

64

Figure 3.8 Effect of relative abrasive hardness on wear rate (Hutchings, 1992)

3.4.3.2 Effect of particle shape

While the effect of particle shape on wear has been researched, only limited test work has

been completed using a quantitative measure of particle shape.

Moore (1980) has looked at the effect of particle shape in abrasion tests on 2 different

white irons. The behaviour of the materials under the same abrasive was markedly different

and is shown in Table 3.8.

65

Table 3.8 Effect of particle shape on relative abrasive wear (after Moore, 1980)

Abrasive Relative Wear Rate Ni-Hard (680 HV)

Relative Wear Rate 28Cr Iron (700HV)

Sharp Crushed Quartz

0.18 0.04

Round Silica Sand 0.03 0.02

Particle analysis using digital imaging techniques has been used by a number of researchers

to determine particle shape. Roylance and Raadnui (1994) use Fourier series analysis of

the perimeter data. They first determine particle size, aspect ratio and then higher order

Fourier coefficients. Stachowiak (2000) has more recently reported on work to measure

particle angularity. The technique called spike parameter quadratic fit (or SPQ)

builds on earlier research that determined a linear fit spike parameter. This was obtained

by fitting differently scaled triangles to perimeter asperities. In the erosion of glass,

Stachowiak found that the SPQ provided fair correlation with the air erosion rate. The

normalised average wear appears to have a linear relationship with the Spike Parameter as

shown in Figure 3.9.

Figure 3.9 Relationship between particle angularity (SPQ) and erosive wear

(Stachowiak, 2000).

66

In the tests, the quartz particles exhibited lower than expected wear rates (relative to their

angularity). This was attributed to the lower fracture toughness of the particles and their

consequent shattering on impact. The shattering absorbs significant collision kinetic energy.

3.4.3.3 Effect of particle size

Particle size is a key determinant of wear severity (down to a point). Mutton (1988)

argues that for metals, particle size has minimal effect at sizes less than 100mm. This is due

to the change in nature of the contact from plastic (d>100mm) to elastic (d

67

Stack and Pungwiwat (1999) found peak slurry wear rate exists using SiC and Alumina

particles in a submerged jet wear tester. For cast iron, peak wear occurred at around

700mm particle size for both erodents. With ceramic materials the peak wear was much

higher (Figure 3.11). It was suggested that the peak wear and the particle size exponent

vary according to the predominant erosion mechanism on the surface. The measured

particle size exponents found in this work (assuming a power law relationship) varied from

1.9-2.4 for Cast Iron to 1.7-2.2 for Alumina ceramic.

Figure 3.11 Effect of particle size on wear rate in slurry jet tester at different velocities

(a), (b) cast iron; (I), (j) alumina ceramic (Stack and Pungwiwat, 1999)

In a Coriolis (sliding bed) wear tester Pagalthivarthi and Helmly (1992) examined the effect

of particle size, ranging from 270mm to 1700mm sand. Velocity and concentration (12 %

Cv) were held constant. The wear rates were seen to increase with particle size up to

about 1300 mm where they appeared to plateau (see Figure 3.12).

68

Figure 3.12 Effect of d50 particle size on wear rate of white irons in Coriolis tester

(Pagalthivarthi and Helmly, 1992).

Apart from the size effect on wear, one of the major issues is choosing a representative

particle size dimension from what is normally a very broad range of sizes in most mineral

processing applications. Roco and Minani (1989) defined an equivalent wear diameter

of a broad sized particle size distribution as the d80 (for pipes) or d85 (for pumps). The

d85 is the sieve or screen size that passes 85% of the particles in the sample. The

equivalent wear diameter is different from the d50 (average particle diameter) because the

interaction energy and rate of material removal are non-linear with particle size (viz. particle

mass and momentum varies as the cube of the diameter).

The equivalent wear diameter d85 will be used to characterise the particle size for slurries

in the current work.

3.4.4 Effect of contact conditions on wear

3.4.4.1 Effect of particle velocity

The effect of particle velocity on wear behaviour is a function of the amount of kinetic

energy dissipated during the collision. Mutton (1988) believes that at velocities less than 10

69

m/s the effect of velocity is not of great importance. However, above this value, and in

particular for pumps, Mutton says the velocity is the most important single factor in

determining wear rate.

The relationship between erosive wear rate (WR) and velocity (V) is commonly considered

to be a power law relationship, WR = k . Vn. The value of the exponent n according to

Mutton (1988) depends on the properties of both the particle and material and the angle of

impingement, but is generally in the range 2-3. The higher n value is associated with softer,

smaller particles, brittle materials and higher impingement angles.

Using the Coriolis (sliding bed) wear tester Pagalthivarthi and Helmly (1992) found that the

wear rate increased approximately with the square of the speed, as shown in Figure 3.13.

Figure 3.13 Effect of velocity on wear rate of white irons in Coriolis tester

(Pagalthivarthi and Helmly, 1992)

3.4.4.2 Effect of impingement angle

Brittle materials such as white irons suffer a change in dominant wear mechanism at higher

particle impingement angles (600 900). The increase in wear rate that results is due to an

increase in the proportion of fragmentation and spalling wear mechanisms (Mutton, 1988).

70

Clark (2001) recently considered the variation in wear rate with impingement angle for

aluminium in a slurry pot test with 550mm silicon carbide particles. The total wear was

proposed to be the sum of the deformation and cutting wear. The total wear peak

occurred at an impingement angle of around 30-400 (Figure 3.14).

Figure 3.14 Effect of impingement angle on wear rate of Aluminium in a pot tester

(Clark, 2001)

3.4.4.3 Effect of concentration

Pagalthivarthi and Helmly (1992) using the Coriolis tester found a linear relationship

between the concentration and wear rate over a range of between 6 and 19% Cv (Figure

3.15). This relationship was explained on the basis that the interaction energy of the

particles depends on the mass flux that, in turn, depends on the concentration.

71

Figure 3.15 Effect of solids concentration (by volume) on wear rate of white irons in Coriolis

tester (Pagalthivarthi and Helmly, 1992)

3.4.5 Erosion and Cavitation

In some applications erosion wear is further complicated by the presence of cavitation

phenomena. Pogodaev and Tsvetkov (1994) state that in sand dredging operations, local

wear is some 8 to 17 times greater than the normal global wear, but that in the presence of

local cavitation the wear factor may reach as high as 22.

Wilson et. al (1992) describe cavitation damage as having an appearance like severe

localised corrosion, with sharp edges being characteristic of cavitation damage. This

appearance is attributed to the high local pressures caused by cavitation.

3.4.6 Erosion and Corrosion

Erosion and corrosion occur together in slurry pumps where metal parts are used in the

presence of a corrosive liquid. The corrosion effects can be significant depending on the

nature of the liquid medium of the slurry and the relative rate of the erosion. Macroscopic

72

surface roughness is a common sign that corrosion and erosion are both occurring

simultaneously (Roco, 1990). The corrosion and erosion processes are recognised as

synergistic and may involve the removal of a passive layer of oxide film which would

protect the surface in an otherwise non erosion environment (Nesic and Postlethwaite,

1981).

Fan et al. (1995) investigated the effect of a typical 26%P2O5 slurry on a high nickel alloy

impeller and found that the local wear was dependant on a critical flow velocity. Once the

velocity was greater than about 12 m/s the erosion rate increased dramatically. This effect

was attributed to the breakdown and removal of the passive film that, in turn, allowed the

corrosion process to accelerate. Preferential corrosion was particularly apparent at the

phase boundaries that appeared sunken in the surface leaving the remaining matrix surface

exposed to the flow.

Other applications where erosion and corrosion can be an issue is where the slurry liquid is

hypersaline or seawater (Shook and Roco, 1991).

3.5 Elastomer wear characteristics

3.5.1 Elastomer mechanisms of wear

The size and shape of the particles and the applied stress levels are important to the wear

mechanisms that occur with elastomer. In addition, the wear of an elastomer is particularly

sensitive to particle impingement angle as this effects the predominant wear mechanism.

Mutton (1988) describes how at low impingement angles tear resistance of the material is

critical due to the process of tensile tearing. At high angles, closer to 900, the resilience is

the important physical characteristic because the collision process is generally purely elastic.

Wear mechanisms such as tensile tearing and fatigue become important with increasing

elastic modulus.

73

Lancaster (1990) describes a number of abrasion wear mechanisms for elastomers that

may also be relevant for erosion processes. In deformation wear, elongation of the material

at the rear of an individual particle contact is followed by tearing and recovery of the

remaining material into a lip. This causes the formation of rows of ridges transverse to the

direction of slurry motion. The tearing process during successive contacts may be

modelled as a propagating crack and thus the wear process becomes amenable to

modelling using fatigue and fracture mechanics concepts. Relating the observed wear with

the elastomers physical and mechanical properties has only had limited applicability due to

the degradation which occurs in the surface layers of the elastomer. The processes that

occur include chain-scission and free-radical reactions. In this case it is the wear process

that leads to the surface layers having different properties from the bulk of the material.

Stachowiak and Batchelor (2001) have described chemical degradation of the surface of

an elastomer. Chemical degradation occurs when hydrophilic particles (such as sand and

silica) impact the surface and the water or oxygen molecules react with the rubber.

Chemical reaction is facilitated by the temperature rise due to repeated particle impact.

The surface layers of the material become mechanically weak and may crack. This process

is illustrated in Figure 3.16.

74

Figure 3.16 Chemical degradation of surface of natural rubber under particle impact

(Stachowiak and Batchelor, 2001)

3.5.2 Effect of elastomer physical properties on

wear

For elastomer materials Mutton (1988) states that the major physical parameters that affect

wear are:

- Elastic modulus

- Resilience

- Friction coefficient

- Tensile strength

- Tear resistance

- Elongation at break

- Hardness

75

Hutchings (1992) notes that attempts to correlate wear rate with the tensile strength of

rubber have been generally unsuccessful. However, high resilience has been found to be

associated with good erosion resistance in a range of unfilled rubber compounds. In some

wear situations a low elastic modulus also leads to a low erosion rate.

3.5.3 Effect of particle properties on wear

No specific test work examining the effect of particle size or shape on elastomer wear was

found in the literature.

3.5.4 Effect of contact conditions on wear

3.5.4.1 Effect of velocity

Lansdown and Price (1986) have described the different wear rate characteristics of a

range of materials. Rubber exhibited a linear increase in wear rate with increasing velocity

up until a critical (unspecified) value. After this the wear rate increased dramatically.

Mutton (1988) recommended a maximum impingement velocity of around 10 m/s for softer

rubbers. The increased wear at high velocities was believed to be a result of scouring or

cutting mechanisms as the rubber is not able to behave elastically.

Hutchings (1992) considered the effect of velocity on wear rate for a natural rubber

undergoing an air erosion test. Wear increased dramatically above about 50m/s as shown

in Figure 3.17. For this material the erosion at 900 impingement angle is around a quarter

the magnitude of that at 300.

76

Figure 3.17 Effect of velocity on wear rate of rubber

(Hutchings, 1992)

3.5.4.2 Effect of impingement angle

At low impingement angles (100 300), Mutton (1988) states that elastomer may suffer an

increase in the tearing and cutting wear mechanisms that lead to increased wear rates.

Conversely, elastomer appears to perform well at higher impingement angles where impact

energy can be dissipated in elastic deformation.

3.5.4.3 Effect of concentration

No specific test work examining the effect of solids concentration on elastomer wear was

found in the literature.

3.5.5 Erosion and depolymerisation

77

Natural rubber undergoes an internal heating on extension and cooling on retraction

(Treloar, 1958). The net energy difference is sufficient at high elongation to cause the

internal and surface temperatures of the rubber to increase. The temperature rise and

hysterisis effect is illustrated in Figure 3.18. As mentioned at the outset of this sub-section,

the effect of multiple particle impact over a prolonged period causes local compression and

a resultant temperature rise that may result in depolymerisation of the material.

Figure 3.18 Effect of elongation on the temperature rise of natural rubber

(Treloar, 1958)

78

3.6 Slurry erosion testing

3.6.1 Introduction

While there is wide understanding of many of the general wear mechanisms for white irons

and to a lesser extent elastomers, it is apparent from the literature that there is only limited

information on specific wear characteristics. Table 3.9 shows a summary of the different

specific characteristics as determined from the literature.

Table 3.9 Specific wear characteristics summarised from the literature

Particle Characteristics Contact Conditions

Relative

hardnes

s

Shape Size Velocity Impingemen

t angle

Concentratio

n

White Irons He

79

conditions chosen were to be reasonably representative of direct impingement erosion wear

inside the pump.

3.6.2 Typical pump operating conditions

In order to choose the range of wear test parameters it was necessary to first understand

the range of typical operating conditions for slurry pumps subject to high wear in mineral

processing plants. These will be reviewed in the context of the Erosion Wear Model

outlined in Figure 3.3.

3.6.2.1 Fluid characteristics

The fluid used in the bulk of mineral processing slurries is water. However, there are a

significant number of applications where either sea water or hyper-saline bore water is

used. While the rheological differences are minor there is an increase in the fluid density

(from 1000 to 1050kg/m3) which will have a minor effect on the buoyancy of particles. A

similar situation exists in the alumina industry where sodium hydroxide (caustic soda) at

concentrations of 15% or more is used. This results in liquid densities of around

1240kg/m3.

For practical test purposes, water has been used as the transport fluid.

3.6.2.2 Particle characteristics

In the bulk of mineral processing applications the particles that are first handled in slurry

form come from the grinding mill. The type of mill determines the ultimate size and

distribution, although maximum sizes of 15mm and average sizes of 500mm are not

uncommon. As the particles are classified in the plant the larger particles are recycled

through the grinding circuit, while the finer particles are subject to further processing. In the

downstream areas and through to tailings disposal the particle size may be only 150mm

80

maximum with 50mm average. A range of typical particle sizes obtained from two different

milling processes is shown in Figure 3.19.

Figure 3.19 Typical mill discharge particle sizes (Kelly and Spottiswood, 1982)

To cover a realistic range of slurry applications, the wear testing d85 particle sizes chosen

were within the range 100-1000mm.

Particle density should be in the range of 2600-3400kg/m3 to cover most of the

applications with silicious ores. This then excludes coal (rs=1400kg/m3) and iron ore

applications (rs=4900kg/m3), although there is a practical benefit in that sand particles

(rs=2650kg/m3) are readily available for testing purposes.

3.6.2.3 Contact conditions

81

To determine the typical contact conditions, a simple analysis was carried out on the inlet

and outlet velocity triangles in a typical 8/6 AH pump operating at best efficiency point flow

and at impeller tip velocities varying from 10 to 35m/s. In the authors experience average

plant duties require impeller tip speeds in the range 20-25m/s. The resultant analysis is

illustrated in Figure 3.20. At the leading edge of the impeller, flow velocities vary from

about 4 to 15m/s (based on the 10 to 35m/s tip velocities). At the trailing edge of the

blade, flow velocities relative to the blade will only be about 2 m/s in the radial direction.

Particle velocities at the tip may be higher than this due to the high coriolis component (ie.

rotating centrifugal field with tangential velocity of 35m/s). At the leading edge the

impingement angles are likely to be high (close to 900), while at the trailing edge on the

pressure side, the impingement angles will be much lower and concentrations higher (due to

the Coriolis forces as mentioned previously).

(a) vane inlet (b) vane outlet

Figure 3.20 Liquid velocity triangles for 8/6AH pump (Q=Qbep; Vt=35m/s)

From the above analysis, flow velocities onto the components would, in the worst case, be

up to 15m/s. Average relative velocities might be around half this. Particle impingement

angles will vary over a range from sliding (ie. zero) to normal to the surface (900

impingement). In-situ particle concentrations will vary with the overall slurry concentration

where particles are directly impinging (ie. without interference). However, concentrations

approaching maximum packing density are possible in areas of highly curved flow, due to

the centrifugal and Coriolis forces discussed previously.

82

3.6.3 Background review of erosion testing

In order to choose the optimum wear test for the above conditions, a review of different

test methods was undertaken.

3.6.3.1 Different erosion testers

Roco et al. (1984) have performed extensive work in both measuring and modelling

erosive wear of centrifugal pumps. In order to establish empirical coefficients for use in

their computational wear model, small-scale experiments were undertaken with a wedge

section specimen inserted in a pipe. This apparatus measured direct impingement wear on

the leading face of the wedge and also simulated sliding-bed type wear along the parallel

trailing edge faces of the specimen. The problem with this type of rig is that it is

cumbersome to use for a large number of different impingement angles and also suffers

from the potential problem of particle degradation due to slurry recirculation. The only test

results reported in this work were for a polyamide resin with impingement angles of 0o and

90o.

A slurry pot tester has been used by Clark (1991) to study the impact of glass spheres on

the surface of copper pin samples. Wear measurements were taken at the stagnation point

of the leading edge of cylindrical specimens rotated in a bath of slurry. Individual particle

impact craters were used to determine impact energies and velocities for the different

particle sizes. Using hard ceramic spheres (SiC) and a relatively soft target material

(copper) avoided particle degradation.

Huggett and Walker (1988) outlined a slurry pot test developed specifically for pump

impeller wear simulation (the disc tester). The immersed fluted disc was designed to

simulate direct and random impingement. The biggest practical problem with this test was

83

again particle degradation. Whilst the overall particle sizing may not change dramatically

during the test, there was high sensitivity to particle shape (particularly particle edge

sharpness). As with most slurry pot testers the disc test does not actually cause break-up

of the particle but just removal and rounding of the sharp edges. Particle shape changes

were found to have affected the wear rate significantly within the first five minutes of

operation. This effect makes the testing of very sharp particles difficult because relatively

small mass loss occurs in the specimen in the short time prior to particle shape changes.

Obviously more rounded particles show less time dependency.

Mens and de Gee (1986) found that when using a two disks slurry pot test that the wear

rate did not change significantly with time. No mention was made of particle shape,

however, and it can only be concluded that the sand used was relatively rounded to start.

The specimens mounted on the contra-rotating discs in this test could be angled to the flow

so that the effects of different impingement angles could be measured. An extensive range

of materials was tested, including elastomers and white irons, with three different sand sizes.

Apart from the slurry degradation problems, the test suffered because of the lack of

knowledge of the flow around the specimen (which influences the accuracy of impingement

angle determination).

A further type of slurry pot test using stationary specimens has been developed by the US

Bureau of Mines (Blickensderfer et al., 1987). This test has the specimens located around

the periphery of a chamber in which a rotating impeller energises the slurry. The specimens

only see low angle (near tangential) wear and the precise angle of impingement cannot be

determined. Degradation problems were overcome by pumping fresh slurry continuously

into the chamber.

To simulate sliding-bed wear Tuzson (1982) developed the Coriolis erosion test that

consisted of a radially oriented spinning pipe attached to a central bowl where the slurry

was introduced. The sample was placed on the back wall of the pipe (in the direction of

rotation). The outlet of the pipe was 1mm square and the rig was limited to using 100mm

84

particles. Pagalthivarthi and Helmly (1992) increased the size of the Tuzson rig so that

larger particle slurries (up to 1700mm) could be tested. This rig can operate on either a

recirculating or once through basis.

Measurement of particle impact velocity cannot be done easily with slurry pot tests because

of the variation in fluid velocity around the specimen. Lynn et al. (1991) used the slurry pot

test of Clarke to complete erosion tests again using silicon carbide particles. Figure 3.21

shows measured impact velocities against the relative free stream velocities.

Figure 3.21 Impact velocity and wear rate as a function of particle size for a slurry pot test

(Lynn et al., 1991)

The considerable decrease in impact velocity for particle sizes 100m and smaller is a result

of much reduced particle inertia. Small particles are more easily constrained by the liquid

flowing about the specimen rather than following an undiverted collision course with the

specimen surface. A further reason for reduced wear rate at smaller particle sizes is the

reduced collision efficiency. This is due to the greater number of particles actually following

the fluid around the specimen. Not only is the impact energy reduced, but the actual

number of collisions is also reduced (for any given solids concentration). The wear rate is

also shown plotted in Figure 3.21. In this case we can see the dramatic increase in wear

with particle sizes in the range of 100-1000m. The larger particles suffer little retardation

85

before impact because of their higher inertia. They also have collision efficiency close to

unity.

A further paper by Wong and Clarke (1993) has looked analytically at the flow around a

cylindrical specimen in the slurry pot erosion tester. Both free stream and boundary layer

flows were modelled. Retardation effects of particles moving in the boundary layer were

found to be critical to the overall wear rate. Close to the specimen surface (within about

half a particle diameter) there is a cushioning effect because of squeeze film separation

and the resultant pressure spike that is generated beneath the impinging particle. Figure

3.22 shows the relationship between free stream velocity and impact velocity for glass

spheres.

Figure 3.22 Impact velocity as a function of free stream velocity for the slurry pot test

(Wong and Clark, 1993)

Based on the experimental results it was found that particles below 100m in size tended to

be trapped in the surface layer and thus the type of wear changed from impinging erosion to

sliding bed erosion.

86

The effect of different impingement angles on slurry erosion rate has been studied by Lin

and Shao (1991). In this work, the specimen was fixed to a rotating arm in a vacuum

chamber, spun and hit by a slurry jet emanating from the ceiling of the chamber. The

orientation of the specimen on the arm could be changed to vary impingement angle. A

range of materials were tested with maximum wear rates for steel being at around 30o

impingement angle moving up to 80o angle for glass. The maximum wear peak angle

decreased slightly with increasing particle size. For angles less than about 70o, particle

impingement angles were about the same or slightly smaller than the liquid impingement

angle. At fluid impingement angles approaching 90o, because the bulk of the solids are in

fact impinging at much lower angles, the effect is to skew the wear rate curve compared to,

for instance, the curve which would be obtained from particles in a gas flow.

Even when using jet type erosion tests, it is important to understand that particle and fluid

trajectories are different depending on particle size. Benchaita et al. (1983) showed that

for particle sizes greater than 2,000m, the particle will impact the material at an angle

equal to that of the liquid jet, but that for much smaller sizes (less than 200m), particles will

follow the fluid and will be mostly swept away. A typical erosion profile for a jet angle of

90o with 700m particle is shown in Figure 3.23.

Figure 3.23 Variation of erosion depth for jet test sample

(Benchaita et al., 1983)

87

The W-shaped profile shows least wear at the stagnation point increasing to a maximum

depth at some larger diameter (which is dependent on the particle size). The greater wear

away from the stagnation point is because of lower impingement angle and higher velocity.

Turenne et al. (1989) studied the effect of particle concentration on overall wear rate in a

slurry jet tester. It was found that erosion efficiency is hindered at high particle

concentrations. The effects of particle screening were greatest for concentrations of more

than 10% by weight. At these concentrations, erosion efficiency remained relatively

constant. The protection effects for higher concentration slurries were noted to be minimal

at around 300 incidence angle because at such low incidence angles there is relatively little

interference from rebounding particles.

Only one of the papers reviewed has undertaken any comparison of both white iron and

elastomer materials. Although Mens and deGee (1986) looked at both white irons and

rubber, they only used rounded particles. As mentioned previously in mineral processing

situations, however, it is the sharp milled or crushed particles which cause the most wear

and which are of greatest interest.

The background review of Section 2 pointed out that in many cases the local gouging wear

is much more important than the global wear. Given this, it is proposed that the major wear

mode involved in local gouging (resulting from a vortex or flow separation) is direct

impingement or random impingement, not sliding-bed. As the Coriolis tester only models

the sliding bed wear mode, use of this type of unit was rejected for this work. Slurry pot

testers were rejected because of the significant particle attrition problems.

For the current work it was decided to use a once through jet type test. This test avoids

the attrition problems of recirculating rigs and has the most flexibility for controlling

impingement erosion conditions, particularly with larger particle sizes (say d85>300mm).

The limitations of the jet erosion test in determining particle angles, as distinct from flow

impingement angles, should be less than for the slurry pot tests. It was expected that test

88

results in the region of 20o - 70o impingement angle and with particle sizes greater than

d85=300mm would be most accurate. Under these conditions inertia forces would

predominate. Overall it was thought that the jet test would offer the least limitations of the

test methods reviewed, allowing independent variation of particle shape, particle size and

impact angle.

3.6.3.2 Comparison of erosion tests and pump wear

Pagalthivarthi and Helmly (1992) examined a number of different test devices and

compared the results with tests on a slurry pump. They found that in direct impingement

wear, over prediction by a factor of 5 was possible when comparing wear at the cutwater.

In other equipment wear simulation, erosion tests on flat plates using a sandblast apparatus

failed to prove a good indicator for wear in pipe bends in extensive tests carried out by

Mills and Mason (1981). Changes in particle velocity and rotation resulting from impact

and fragmentation were the main reasons the experimental erosion tests were felt to be not

more representative (Finnie, 1995).

Wilson et al. (1992) believe that the major difficulty in erosion wear testing is combining the

impingement and sliding wear tests back into a result which is representative for the pump.

This is an important point as they note that there appears to be different scaling factors for

the two different wear types. They also state that the only way to be sure that the results

are accurate is to verify the test with wear measurement on a full-scale pump!

Using the jet test to determine material characteristics may prove a more reliable indicator

for pump wear in the field. As the wear will be measured over a range of impingement

angles it may be possible to better correlate field wear in the pump at a specific location

with a specific impingement angle from the test.

3.6.4 Experimental equipment and method

89

The jet test used in this work is basically an eductor where the solid particles are entrained

in a liquid stream and blasted onto the surface of the sample. The particles are loaded

through a funnel from a hopper above the eductor and high-pressure water from a positive

displacement pump is fed through a nozzle into the mix-chamber. This slurry is then forced

through a ceramic nozzle of 4.5mm diameter onto the test specimen, which is placed 20mm

from the end of the nozzle. The apparatus is shown in detail in Figure 3.24.

(a) Funnel feed and water inlet pipe (b) Sample holder and nozzle

(c) Schematic of equipment

a

90

Figure 3.24 Details of slurry erosion jet wear tester

The sample holder can rotate to give a variety of impingement angles (a) between 150 and

900. The used slurry is collected in a waste hopper underneath. The speed of the positive

displacement pump could be varied to alter the water flow rate. The slurry jet velocity

could be easily varied between 10 and 20m/s. Both the solid particle sizing and the liquid

velocity determine the concentration of the slurry in the nozzle.

Two different solids were used in the test work to simulate different particle shapes. For the

rounded particles, four different quartz river sand particle size distributions with

d85=150mm, 350, 500mm and 1000mm were chosen. For sharp particle applications,

fused alumina with d85=100mm, 550mm and 950mm were selected. Individual particle

sieve analyses are shown in Table 3.10.

Table 3.10 Particle size distributions (a) sand

Cumulative % passing Sieve Size

(mm) d85=150mm d85=350mm d85=500mm d85=1000mm

53 10

75 30

106 65

150 85 1

180 93 0.2 3

250 98 21 15

300 100 63 27

425 100 79

500 96

600 100 3

850 25

1000 99

1180 100

91

92

(b) fused alumina

Cumulative % passing Sieve Size

(mm) d85=100mm d85=550mm d85=950mm

53 3

75 24

106 92

150 99

180 100 1

250 4

300 6

425 30

500 66

600 95 2

850 100 14

1000 97

1180 100

Particle shape analysis was undertaken by the University of WA using the method of

Stachowiak (2000), with the results shown in Table 3.11(a) and (b) for the sand and

alumina materials respectively.

Table 3.11 Particle shape factor (SPQ)

(a) sand Nominal d85

(mm)

Max. dimension

(mm)

Number of

particles

measured

SPQ Standard

Deviation

150 80-200 25 0.40 0.10

500 230-890 25 0.31 0.13

1000 950-1890 25 0.33 0.11

(b) alumina

Nominal particle range

(mm)

Number of particles

measured

SPQ Standard

Deviation

100-500 25 0.53 0.16

500-1000 12 0.42 0.12

93

Photographs of the particles are shown in Figure 3.25. The fine particles were in general

sharper than the larger sizes with a SPQ some 20-25% greater. The alumina particles were

sharper than the sand with SPQ 25% larger.

(a) sand 150mm (b) sand 1000mm

(c) alumina 100-500mm (d) alumina 500-1000mm

Figure 3.25 Sand and alumina particle shape

The materials tested in the jet tester included (by material code) white irons A05 and A12,

rubbers R08, R26 and R33 and polyurethane U01. Descriptions of these materials have

been given previously in Section 3.3.

A standard test piece of the material was either moulded in the case of polymers or, in the

case of metal, cut from a large cast test block. The surface of the metal test samples was

ground, but the other parts were used as moulded. The 40x50x10mm specimens were

94

weighed and secured into the sample holder prior to testing. Slurry impingement angle was

set using a graduated scale on the side of the test chamber. A total of between 10 and

15kg of wear media was generally used and the time taken for the wear media to pass

through the eductor mix chamber was recorded in order to calculate the concentration of

solids in the slurry jet. The test period varied but was in the order of 5-10 minutes. After

testing, the specimens were dried for 24 hours, air cleaned and weighed to 0.0001g. The

wear rate for the test piece was then calculated as volume loss per kilogram of wear media

blasted at the sample (ie. mm3/kg).

3.6.5 Results

The number of tests undertaken was quite extensive, with over 50 individual tests

completed on each of the materials. The full data set of all the test results is included at

Appendix A1.

Slurry concentrations were found to vary between 6% and 14% (by weight) over the range

of jet velocities. This should have ensured a good collision efficiency and minimum

interparticle interference.

Based on the preliminary velocity analysis in the impeller, the original intention was to

operate with jet velocities in the range 10-15m/s. It was found that at 10m/s jet velocity,

only the large alumina particles wore sufficiently to achieve any measurable weight loss.

Because of this, the test velocity range maximum was increased to 20m/s.

The small sand particles (d85=150mm) caused such low wear rates that they were

immeasurable for all the samples. The elastomer in particular simply showed no change in

surface appearance after the test run. For the larger sand particle sizes (d85=500 and

1000mm) the rubber samples showed evidence of surface de-polymerisation (sticky and

hot - up to 800C) at the 20m/s jet velocity. As a result, most of the sand testing had to be

undertaken at the 15m/s velocity, with the 20m/s results limited to the alumina particles.

95

The profile of the wear gouge varied with impingement angle. Even at 900, however, there

was no evidence of the W-shaped wear profile reported by other researchers. Most of

the grooves showed a uniform concave shape as shown in Figure 3.26. The difference in

shape seen with the current test work may be a result of the relatively small jet diameter

and relatively large particle sizes (smallest ratio 4:1).

(a) A05 (b) R08

Figure 3.26 Samples showing wear groove after testing

(Vj=20m/s; a=200; d85=950mm alumina)

3.6.5.1 White iron wear mechanisms

Due to the particle size and velocity of impact, the major wear mechanisms for the white

irons were fracturing and ploughing. Figure 3.27 shows the surface of the A05 and A12

samples at the perimeter of the wear gouge with a number of clearly defined ploughs marks

caused by individual particles. In Figure 3.27(b) a number of carbides can be seen at the

surface standing out from the surface. This as though they have been fractured away

from the matrix by the particle moving across it.

96

(a) A05 (b) A12

Figure 3.27 White iron sample surface showing individual particle gouges

(Vj=20m/s; a=150; d85=1000mm sand) mag. 270x

Figure 3.28 shows more closely the A05 worn sample from Figure 3.27(a). The individual

carbides are cracked both at the grain boundary and within the carbide itself. This would

ultimately allow sections of carbide to break away on subsequent particle impacts. In this

instance this was the major wear mechanism.

Figure 3.28 A05 surface at higher magnification showing cracked carbides

(Vj=20m/s; a=150; d85=1000mm sand) mag. 1000x

97

3.6.5.2 Elastomer wear mechanisms

Figure 3.29 shows the worn surface of the R26 rubber. Near the perimeter of the wear

gouge the effects of individual particle impacts can be seen. Figure 3.29(a) shows the sand

wear, while (b) shows the alumina particle wear. The cutting with the more rounded sand

particles was similar to the alumina, although the cut width was greater and the tear was not

as deep at the initiation point.

(a) d85=1000mm sand (b) d85=950mm alumina

Figure 3.29 R26 sample surface showing scratches, cuts and tears

(Vj=20m/s; a=70) mag. 70x

Figure 3.30 shows a further magnified image of the surface wear with the 950mm alumina

particles. The cuts, tears and ploughs (where material has been completely removed) are

clearly visible. Also visible is some of the wear debris.

Figure 3.30 R26 surface at higher magnification showing cutting and tears

98

(Vj=20m/s; a=70; d85=950mm alumina) mag. 500x

Figure 3.31 shows the R26 rubber worn sample surface at high angle impingement (a=900)

with the 550mm alumina particles. In Figure 3.31(a) ploughs, gouges and cuts are clearly

visible. At higher magnification, the worn material surface in Figure 3.31(b) and (c) looks

to have suffered some surface breakdown with small plate like or layered structures

apparent, interspersed with relatively deep holes. This may be a form of chemical

breakdown or depolymerisation of the surface layer as described in Sub-section 3.5.5.

(a) mag. 300x (b) mag. 1500x

(c) mag. 3000x

Figure 3.31 R26 worn surface at high magnification showing breakdown of polymer structure at

high jet impingement angles

(Vj=20m/s; a=900; d85=550mm alumina)

99

3.6.5.3 Effect of impingement angle

The effects of impingement angle on wear rate were tested for the two different solids (sand

and alumina).

For the four elastomers the trend in the results is similar. Maximum wear was seen at an

impingement angle of between 20-300. The maximum occurred at the same impingement

angle for both the sand and the alumina. Typical wear results for the R33 material are

shown in Figure 3.32.

Figure 3.32 Effect of impingement angle on R33 material wear (Vj=20m/s; alumina particles)

The different white irons tested showed a similar maximum in the curve to the elastomers

but at a greater angle of around a=30 to 400 impingement (see Figure 3.33). The relative

maximum wear was not as great as the elastomer. The maximum wear rate occurred at the

same impingement angle over the range of particle sizes.

0.00

5.00

10.00

15.00

20.00

25.00

0 20 40 60 80 100

Impingement angle (deg)

Wea

r ra

te (

mm

3/kg

)

d85=100umd85=550umd85=950um

100

Figure 3.33 Effect of impingement angle on A05 wear (Vj=20m/s; sand particles)

As a result of the identification of the maximum wear rate values at around a=300 for both

material classes the bulk of the latter tests were restricted to this value and also the highest

impingement angle of a=900.

3.6.5.4 Effect of particle shape

The wear rate for varying particle shape is shown for the white irons at impingement angles

of a=300 and 900 in Figure 3.34. The effect of shape on wear rate as measured by the

SPQ was different for the different particle sizes. At the smaller sizes there was less effect

than at the largest size. Overall, the effect of particle size on wear was similar for both A05

and A12.

On the basis that wear rate is low for small SPQ values (eg. round particles), a power law

relationship has been used to fit the data. Power law exponents for the different particle

sizes at a=300 varied from 5.8 for d85=300mm down to 2.4 for d85=1000mm. At the

0.00

1.00

2.00

3.00

4.00

5.00

6.00

7.00

8.00

9.00

10.00

0 20 40 60 80 100

Impingement angle (deg)

Wea

r ra

te (

mm

3/kg

)

d85=300um

d85=500um

d85=1000um

101

higher impingement angle shown in Figure 3.34(b) the exponents varied randomly with size

within the range 1.7-3.7.

(a) a=300

(b) a=900

Figure 3.34 Effect of particle shape factor on white iron wear rate at different impingement

angle (Vj=20m/s)

Figure 3.35 shows the effect of particle shape for two of the elastomers (R08 and U01).

At a=300 impingement angle the power law exponent varied from 2.1-5 for the R08. For

the two different particle sizes for the U01 material the exponent was the same value of 3.1.

No a=900 impingement angle results were generated for the elastomer materials because

0

5

10

15

20

25

0 0.2 0.4 0.6

Shape Factor SPQ

Wea

r ra

te (

mm

3/kg

) A05: d85=300um

A05: d85=500um

A05: d85=1000um

A12: d85=300um

A12: d85=500um

A12: d85=1000um

02

468

1012

141618

0 0.2 0.4 0.6

Shape Factor SPQ

Wea

r ra

te (

mm

3/kg

)

A05: d85=500um

A05: d85=1000um

A12: d85=500um

A12: d85=1000um

102

of the very low volumes of material removed.

Figure 3.35 Effect of particle shape factor on elastomer wear rate (Vj=20m/s; a=300)

3.6.5.5 Effect of particle size

To eliminate the effects of particle shape and thus be able to examine the size effect only, all

the wear results were corrected to SPQ=0.31 (ie. the shape factor at d85=500mm) using a

power law relationship with exponent 3.

Figure 3.36 shows the effects of particle size on material wear for both impingement angles

of 300 and 900 for both the sand and alumina slurries. The wear rate can be seen to

increase with increasing particle size, following approximately a power law relationship.

Reasonable correlation can be seen between the sand and alumina results (as there should

be with the shape factor correction applied). There were some outlying points for the

d85=300mm particles. The a=900 impingement results appear to follow the same

relationship with particle size as the a=300 results although at a lower relative wear rate.

The wear rate exponent varied considerably with the polyurethane value of 1.4, the white

iron value between 1.2-1.9 and the rubber value 1.1.

0

5

10

15

20

25

30

35

40

0 0.2 0.4 0.6

Shape Factor SPQ

Wea

r ra

te (

mm

3/kg

)

R08: d85=100um

R08: d85=500um

R08: d85=1000um

U01: d85=100um

U01: d85=500um

U01: d85=1000um

103

(a) white iron

(b) elastomer

Figure 3.36 Effect of particle size on wear rate for different materials (a=300, 900; Vj=20m/s)

Figure 3.37 shows all of the a=300 impingement results plotted on the one graph. In

general, the results show that the rubber wear rate was less than the white irons. However,

in all cases the polyurethane (U01) wear rate was the worst.

0.1

1

10

100

10 100 1000 10000

d85 (micron)

Cor

r. w

ear

rate

(m

m3/

kg)

A05 30degA12 30degA05 90degA12 90deg

0.1

1

10

100

10 100 1000 10000

d85 (micron)

Cor

r. w

ear

rate

(m

m3/

kg) R08 30deg

R26 30deg

R33 30deg

U01 30deg

R08 90deg

R26 90deg

R33 90deg

104

Figure 3.37 Effect of particle size on wear rate for all materials (a=300; Vj=20m/s)

3.6.5.6 Effect of velocity

Wear rate was measured at three different jet velocities; 10, 15 and 20m/s. Only the 15

and 20m/s results were used because the wear rate was too low and inconsistent at the

lower value (as mentioned previously).

The wear rate for different jet velocities is plotted for the white iron parts in Figure 3.38.

Based on the previous literature, a power law relationship was fitted to the data giving

exponents ranging from -0.4 to 2.0. Whether it is just scatter or a real trend, the difference

in slope of the A12 and A05 results indicates a cross-over if the data is extrapolated to

lower velocities. This would result in the A12 wear rate being lower than the A05. This is

a not unlikely result given the change in particle impact energy and the beneficial effect that

will have on the large carbide fracturing of the A12.

0.1

1

10

100

10 100 1000 10000

d85 (micron)

Cor

r. w

ear

rate

(m

m3/

kg)

R08

R26

R33

U01

A05

A12

105

There is also a cross-over between the alumina particle results for the A05 material. The

d85=950mm wear rate being lower than the d85=550mm result. This is likely to be

experimental error and is not supported by the other results shown in sub-Section 3.6.5.5

for the effect of particle size on wear rate.

Figure 3.38 Effect of jet velocity on wear rate for the white irons (a=300)

For the rubbers (shown in Figure 3.39), the wear-velocity exponent was much greater than

the white irons, in this case averaging at about 2.4 for the rubber and 4.7 for the

polyurethane. The high sensitivity of the elastomer wear rate to impingement velocity has

significant implications for extrapolation to conditions outside the test range. Whilst the

overall wear rates across all the materials for velocities in the range 15-20m/s are not too

dissimilar, the relative performance of the white irons and elastomers would differ

dramatically at velocities below about 10m/s.

02468

101214

1618

20

0 10 20

Jet velocity (m/s)

Wea

r rat

e (m

m3/

kg)

A05; d85=1000um sand

A05; d85=500um sand

A05; d85=550um alumina

A05; d85=950um alumina

A12; d85=500um sand

A12; d85=1000um sand

A12; d85=550um alumina

106

(a) Rubber results

(b) Polyurethane results

Figure 3.39 Effect of jet velocity on wear rate for the elastomers (a=300)

02468

10121416182022

0 10 20Jet velocity (m/s)

Wea

r rat

e (m

m3/

kg)

R08; d85=1000um sand

R08; d85=550um alumina

R08; d85=950um alumina

R33; d85=1000um sand

R33; d85=550um alumina

R33; d85=950um alumina

R26; d85=1000um sand

R26; d85=550um alumina

0

10

20

30

40

0 10 20Jet velocity (m/s)

Wea

r rat

e (m

m3/

kg)

U01; d85=500um sand

U01; d85=1000um sand

U01; d85=550um alumina

U01; d85=950um alumina

107

3.6.6 Discussion and analysis

3.6.6.1 Wear mechanisms

The white iron wear mechanisms found in the tests were similar to that reported in the

literature. The particle impacts form individual plough and scratching marks as described

by Finnie (1995). For the larger carbides of the hypereutectic white iron A12, it is

apparent that the carbides are fracturing, both within the carbide and where subject to high

impact, at the carbide boundary as well. The effect of the different phase hardness is

visible with preferential scratching occurring in the matrix relative to the harder carbide.

Certainly the finer carbide structure of the A05 did not offer the same opportunity for

carbide fracturing and pull-out observed for the A12.

The wear mechanisms for the elastomer also followed (at least at the macro level) the

descriptions of the literature, in that impacting particles gouged relatively large chunks of

material or sliced or cut the surface. At high magnifications it was interesting to observe an

entirely different wear mechanism with small plate like structures, and relatively deep

fissures being visible. This may be part of the chain-scission process that has been

mentioned but not described in detail in the literature by Lancaster (1990). Polymer

breakdown as part of the wear process is not well understood and is beyond the scope of

the project here.

3.6.6.2 Effect of particle properties

The measured effect of particle shape on wear rate in slurry erosion has been investigated

for the first time in the work here. The effect of particle shape on wear was significant for

both white iron and elastomer material classes. Wear rates increased with an increase in

the sharpness of the particle. For the white irons, this is probably the result of the higher

contact pressures of the sharp particle asperities which cause either carbide fracturing or

108

ploughing/scratching of the matrix. For the elastomer material, the sharper particles will

more easily cut the surface which in turn provides an initiating point for fatigue crack

propagation or tearing.

Whilst there was some scatter in the measured SPQ values, overall correlation with the

wear rate using a power law exponent of 3, was quite good. This is different to the linear

relationship established by Stachowiak (1998) in the air erosion tests. Reasons for the

different relationship are not readily explainable. The number of data sets for both the

Stachowiak air erosion tests (4 different particles) and the slurry erosion tests undertaken

here (2 different particles) are very limited and further research is required.

For the white irons there was an increase in the SPQ power law exponent for decreasing

particle size, indicating a greater sensitivity to particle shape. This may be a result of the

particle and its asperity size relative to the scale of the carbide microstructure. Smaller

particle protuberances would be able to remove matrix material between carbides, whereas

the matrix would be protected from larger particles by the carbides.

The effect of relative hardness of the particle to the material can be discounted as having

had a major effect on the white iron results. As discussed previously in sub section 3.4.3

(Hutchings, 1992), this is only of significance when the He/Hm

109

Table 3.12 Comparison of particle size exponent for different wear tests

Coriolis tester

d

110

a=300 and a=900 is some 2-3 times for the elastomers, but significantly less for the white

irons. The lower wear rate at the high impingement angles can be explained by the

elastomers ability to absorb the particle impact energy under compression. Only with high

energy impacts will the elastomer tensile stress be exceeded and localised material failure

occur. For low particle velocities or small particles, the collision may be totally elastic with

no detrimental effect on the impacted material. This contrasts with the white iron situation

where due to its high relative hardness, high angle impingement will likely cause fracturing of

the carbides leading to high wear rates. This appears to be the same fracturing wear

mechanisms that occurs at lower impingement angles as well.

The absolute size of the particle is important because it influences the impact energy.

However, it is thought that the relative size (ie. the particle size compared to the scale of the

white iron microstructure) also plays a role in determining wear rate at the micro level. The

particle edge asperity scale is believed to be important. If the particle is large, or if the

asperity is small then it will have the wear effect of a small, particle even if the overall size is

very much greater. This is why refining the microstructure of white irons is successful

(Dolman, et. al., 1998). The intercarbide spacing is reduced relative to the particle size.

Another important factor may relate to the effects of fluid drag on the smaller particles. As

explained previously, at low impingement angles fewer particles may be reaching the

material surface due to drag forces exceeding inertial forces. However, at higher

impingement angles, there would be more particles striking the surface due to the tighter

radius of curvature of the flow.

The wear rate as a function of jet velocity was found to vary considerably depending on the

material. Exponent values ranged from 0.8-4.7. This is a different result to the general

literature that shows exponents only in the range of 2-3 only (Hutchings, 1992). The wear

exponent did not appear to be greatly influenced by either particle shape or particle size,

just the material properties. The higher poly and rubber results may be due to the velocity

(or impact energy) exceeding the critical value as described by Lansdown and Price

111

(1986). Certainly the jet velocity is greater than the 10m/s value that Mutton (1988)

proposed as a maximum for soft rubbers. Understanding the velocity relationships properly

requires further investigation and is outside the project scope. The lower wear rate for the

A12 relative to the A05 material at low velocities, as mentioned previously, is probably a

function of the change in wear mechanisms (less carbide fracturing).

3.6.6.4 Empirical wear relationships

Using the general form of the erosion wear rate equation from Finnie (1995) and Hutchings

(1992) and adding terms for particle size and shape factor, a new general empirical form is

proposed in Equation 3.1.

WR = M.K.SPQp.Vjn.d85m.f(a) Equation 3.1

where: M = mass of particles contacting surface, a f(Cv, rs)

K = material constant

p = 2-6 for white irons

= 2-5 for elastomers @ a=300 n = 0.8 for white irons @ a=300

= 2.4 for rubber @ a=300

= 4.7 for polyurethane @ a=300

m = 1.1 for rubber

= 1.2-1.9 for white iron

= 1.4 for polyurethane

This equation may provide a starting point for the estimation of pump wear to be

considered further in Section 8.

3.6.6.5 Material Wear Life Factors

112

A common method of representing the relative performance of materials, especially useful

from the selection point-of-view, is to use a Wear Life Factor (WLF). Although the WLF

will obviously vary with particle size, shape and contact conditions (as indicated above) an

average value still has usefulness for general comparative purposes in applications in the

field.

The WLF here is calculated with respect to the A05 material (eutectic white iron), which is

a well-known and widely used material in slurry pumps. The WLF is defined as:

WLF = (Wear rate reference A05) / (Wear rate of material of interest) Eq. 3.2

WLFs for the tested materials have been extrapolated for Vj=20m/s and 10m/s using

Equation 3.1 above and are shown in Figure 3.40 (data points only shown to differentiate

the materials). This illustrates the large effect that the velocity has on the wear rate of the

elastomer materials in general and polyurethane in particular.

(a) Vj=20m/s

0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

0 200 400 600 800 1000 1200

d85 (micron)

WLF

A05A12U01R08R26R33

0

2

4

6

8

10

12

14

0 200 400 600 800 1000 1200

d85 (micron)

WLF

A05

A12U01R08R26R33

113

(b) Vj=10m/s

Figure 3.40 Extrapolated Wear Life Factors @ SPQ=0.3; a=300

A major anomaly in the WLF calculations is the A12 hypereutectic white iron. This may be

due to the sensitivity to particle impact energy, mentioned previously in the results section.

The field performance of the hypereutectic materials is definitely better than that indicated in

the jet test results above (Warman internal communication, 2000). It is possible that the

high velocities of the jet test are causing fracturing of the carbides to be the main wear

mechanism. This may not be the case at lower impact energies as indicated by

extrapolation of the results back to the 10m/s axis. If this is so, the larger volume of harder

carbides in the A12 would better resist the scratching wear that appears to dominate.

The examination of the field parts wear mechanisms in Sections 6 and 7 should help

determine if this is actually the case.

3.7 Summary and Discussion

This section has reviewed the major literature that relates to slurry erosion wear of white

iron and elastomer. In general, only a small amount of research has been done using

material, particle and contact conditions that are representative of slurry pumping duties in

the mineral processing industry.

The major erosion wear modes for slurry pumps identified in the literature are direct

impingement and sliding bed. Modes such as random impingement are also significant.

The hypothesis proposed here is that local gouging wear (identified in Section 2 as the life-

limiting wear in a pump), is largely a function of the direct impingement of particles. This is

especially the case when those particles are caught in a highly radial flow such as a vortex

or strong eddy and subject