Wear, 84 (1983) 39 - 49 39

PREDICTION OF ABRASION WEAR FOR SLURRY PUMP MATERIALS

AHMED ELKHOLY

Pratt and Whitney Aircraft of Canada Ltd., Mississauga, Ontario L5A 324 (Canada)

(Received June 3, 1982)

Summary

Coal and other mineral pumping installations have been proven to be practical and in most cases economical. However, pump wear is still a highly important factor in the total economics of slurry handling. In this study, a series of tests were undertaken on different pump materials to study the effect of slurry properties and pump material on abrasion wear using a test facility which incorporates most of these parameters. An analytical model for the prediction of wear for hard and brittle materials under different run- ning conditions was obtained. Experimental and calculated results for cast iron as an example of a hard and brittle material are comparable on a speci- men weight loss basis.

1. Introduction

Large-scale transportation of solids by pipeline has moved from the status of a rather risky possibility to its present status of an excellent alterna- tive to the conventional transportation modes. The largest force behind the growing interest in solids pumping has primarily been one of economics. Transportation costs as a function of annual throughput for three different modes of transportation are shown in Fig. 1. They are not greatly affected by the length of the system once it is longer than 50 - 80 km. Also, Fig. 1 shows that there is a larger benefit for pipeline transportation as the annual throughput is increased.

Two factors are common for a successful application of slurry pipelines to compete with other transportation modes, the first factor being that the requirements of volumes are large enough to justify a slurry pipeline. The second is that for a cross-country solids pipeline to be practical, the slurry must be designed to be stable hydraulically. This means the slurry can be shut down in the pipeline and restarted without difficulty. When examining slurry pipelines as an alternative transportation mode, the pumps and pipe- lines which constitute the system will experience extensive wear. This wear phenomenon must be minimized for the system to be efficient and econom-

0043-1648/83/0000-0000/$03.00 @ Elsevier Sequoia/Printed in The Netherlands

1.

.8

A

.4 -1 .2 .4 .6 .8 1. 2. 4. 6. 8. 10.

ANNUAL THROUGHPUT (MILLION TONS)

Fig. 1. Transportation cost8 as a function of annual throughput.

ical. Although much research has been carried out in the field of wear in the last decade, no usable formula seems to have been developed for the predic- tion of slurry abrasion.

The main objective of the study is to develop an analytical model for the prediction of the amount of wear in a slurry pump under specified condi- tions. The approach once the analytical model is developed is to correlate the model with the decrease in pump performance and ultimately to develop methods to increase the pump life. Therefore the effects of certain param- eters on the amount of wear resistance of particular materials used in the pump industry have been investigated.

2. Experimental modelling technique

In experiments involving abrasion resistance it is essential that the test procedure be kept as standardized as possible. This is due to the variety of types of wear that the term abrasion may include.

A large variety of facilities used for abrasion testing have been suggested in the literature [l - 71. The literature also provided the following list of parameters which were thought to play significant roles in slurry pump wear: (1) the velocity V (m s-l) of the particles; (2) the impingement angle LY (deg); (3) the ratio H&I, of the solid particle hardness H2 (HB) to the eroded ma- terial hardness H, (HB); (4) the solid particle size d (mm); (5) the slurry con- centration C, by volume; (6) the difference ,S,--Ss between the specific gravity of the fluid and the density of the particles; (7) the viscosity /.I (kg m-’ s-l) of the liquid phase.

41

d 8

Fig. 2. Erosion wear test rig: 1, test tank; 2, specimen holder; 3, protector plate; 4, sand entrance; 5, norbide nozzle; 6, centrifugal pump; 7, diaphragm pressure gauge; 8, Jaco air pinch valves; 9, air supply; 10, variable speed motor; 11, weight scale; 12, graduated pail; 13, sample nozzle.

The jet abrasion test facility shown in Fig. 2 was designed from the background experimental modelling technique literature to evaluate wear resistance of different materials which are being subjected to abrasive parti- cles. The main asset of the test facility is its simple design although each wear parameter can be measured.

With a variable speed motor slurry velocities ranging from 5 to 30 m s-l at the nozzle exit can be obtained. Measurement of the slurry velocity is readily carried out by instantaneously redirecting the flow with a Jaco air-operated pinch valve through the same nozzle size into a graduated pail. The length of rubber pipe between the T-section to each nozzle is chosen to be equal for consistant simulation. By recording the time of sampling to- gether with the volume of sample collected by reading the level from a sight tube on the side of the pail, the nozzle exit velocity is obtained from

J7=Q _?_ ms-’ TsD2 3n

(1)

42

where Q (1) is the volume of sample collected, Z’, (min) is the time of sam- pling and D (cm) is the diameter of nozzle.

From the sample collected the density y of the mixture is obtained by dividing the mass of sample by the volume of the sample. The concentration C, of solid particles by volume is calculated from

C” = sin-SL

ss-SL (2)

where S, is the specific gravity of the mixture, Ss is the specific gravity of the solids and SL is the specific gravity of liquid. The concentration C, of the solid particles by mass is obtained from

cw=y (3) m

A protector plate is initially located above the specimen holder to allow complete mixing of the sand and water before removal and contact of the specimen by the jet is made.

The particle size of the sand sample is determined by selecting one of four different mesh sizes ranging from 37 to 3; or equivalent openings of 0.41 - 5.6 mm. A particle size range of 0.41 - 0.50 mm was used for experi- mental verification in this study. The impingement angle can be varied from approximately 15O to 90” by adjusting the specimen holder. The volume of the liquid phase is determined from a sight tube calibrated in litres mounted on the side of the testing tank.

A pressure gauge fitted with a stainless steel diaphragm seal is located in the pump discharge piping for the purpose of indicating approximate veloc- ities at the nozzle exit. The seal must be kept away from the moving par- ticles; therefore a tube elbow mount as shown in Fig. 2 was used. The tube has to have a large internal diameter compared with the particle size to avoid any clogging.

The main limitation of the test facility is that only one specimen can be tested at one time. Being a closed loop, particle degradation is evident. The weight loss rate cannot be measured, which introduces another parameter to the investigation.

After the specimen is left under the effect of the slurry jet, weight loss due to wear is observed. The wear index is calculated from the following equation :

w = Q-YCwTw x 103

Ts Wlo ss

(4)

where y (g cme3) is the density of the mixture, T, (min) is the time (given by time of testing - sampling time) of wear and Wloss (g) is the specimen

43

mass loss. The wear index W as given in eqn. (4) is expressed as grams of specimen removed per gram of sand. Wear can be expressed as an absolute value of mass loss or as a wear index as shown in eqn. (4),

3. Effect of different parameters of wear

3. I, Velocity Two different materials, aluminium and cast iron, with Brine11 hard-

nesses of 121 and 230 respectively, have been tested to verify the wear mechanism model. Aluminium was chosen to represent soft and ductile ma- terials while cast iron was chosen to represent hard and brittle materials. The ideal materials would be a medium to high ahoy white cast iron known as Ni- Hard and slurry pump lining elastomers. Because of the limited av~lability in small samples of Ni-Hard and the property of rubber of water absorbance, the aforementioned materials were used for experimental verification,

By maintaining a constant concentration of solid particles in the liquid phase, particle size and impingement angle for each material and varying the velocity, a corresponding value of the wear is generated. After three or four velocities and wear values are obtained, the results are fitted to a curve of the form

W=KVn (5)

where W is the wear terms of the mass in grams of specimen removed per gram of sand, K is a constant term which accounts for all the remaining parameters, V is the velocity of the particles at the nozzle exit and n is an appropriate exponent. An equation for each individual material is obtained and plotted on a log-log scale (Fig. 3) to determine the velocity exponent. A sample of Belzona molecular ceramic steel (BMS) was obtained by mixing a 3:l ratio by volume of base to solidifier and letting it cure for 24 h. BMS is a low temperature solidifying, stress, chemical and heat-resistant molecular metal used for rebuilding machinery and equipment. BMS was tested under identical conditions as for aluminium and cast iron and the results are plotted in Fig. 3. By choosing any arbitrary velocity and comparing the ceramic steel with the other two materials it can be seen that the ceramic steel has a higher wear index and consequen~y a poorer wear resistance.

3.2. Impingement angle Further tests have been conducted on cast iron to break down the con-

stant term by changing the impingement angle to 90’. By assuming that the impingement angle QI is independent of velocity, an equation was developed for cast iron as an example of a hard and brittle material in the form

W=1.061X10V8X CY - or

go-ecu, (6)

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2

1

1 2 5 10 20 50 100 VELOCITY,m/s

Fig. 3. Velocity-wear relationship: 0, aluminium; 0, cast iron; +, BMS.

where ol is the angle at which wear develops. It was taken to be zero for sim- plicity and practical application.

By inserting the experimental velocity values into eqn. (6), correspond- ing wear values are calculated. Power curves are fitted to the calculated wear values at 60’ - 90” impingement angles for a comparison with the experimen- tal results (Fig. 4). The 90’ impingement angle curve shows a higher wear index for any arbitrary velocity value, as expected.

3.3. Hardness The effect of hardness is shown in Fig. 5, where the hardness is ex-

pressed as the ratio of the particle hardness to the eroded material hardness and the wear is expressed as a mass loss. Experiments were performed on four different samples of various hardnesses ranging from 150 to 400 HB. All the parameters affecting the wear were held constant while the particle size was varied and the wear as a mass loss was recorded. The abrading material, which is silica sand, has a hardness of 710 on the Brine11 scale. As can be seen, there are two distinct slopes for each different particle size. When the

\ ‘\ \ \

0 P

: 0 : 0 0 n

2 0 ml

+OTX (INUS JO b&-83 HI3d NLlW133dS JO WV83 'XYL'M

46

hardness effect is considered it is imperative owing to the steep slope below the transition point to know the exact hardness ratio.

From tests conducted on various materials using the same silica sand, it was found that the wear varied with the hardness ratio with an exponent of 3.817 below the transition point and with an exponent of 0.268 above this point.

3.4. Particle size In general the absolute wear increases with increasing particle size. All

the parameters affecting wear were held constant for all the experiments. The impingement angle was 30” and the time of wear was 2 min. The charac- teristic curves of wear versus particle size for various materials are shown in Fig. 6. It was found that the amount of wear varied with the particle size with an exponent of 0.616.

3.5. Concentrations It was found that wear increases with increasing solid particle concen-

tration in slurries. The mass loss of plain carbon steel samples versus volu-

1 ___-__ _ .a ___-

.6

.4 _ __- C”ZlT%

\

ii *2 :ll%‘.

\

1 x -__ I _

m- :7% 1

k? t? P v:

z .l -4 6 -. 0

.08 ___-__ ___-

.06 _ ____.. .- -7 /

.04

.Ol : i : _

.l .2 .4 .6 .a 1 2 4 6 8 IO

PARTICLE Sl%i:, mm

Fig. 6. Particle size-wear relationship (n = 0.616).

1

.6

.6

.02

.Ol

1 2 4 6 6 10 20 40 60 80 100

Cv%

Fig. 7. Concentration-wear relationship.

metric concentration C, is shown in Fig. 7 for various particle sizes. It was found that the amount of wear varied with the concentration with an expo- nent of 0.682.

4. Derivation of the analytical model

By combining all the parameters from the previous section an equation of the following form is produced:

A W = $.KCVo,682 s: 180 - 90 f(t) g (7) 1

where n is a function of H,/H2 according to

n = 3.817 if 3 < 1.9 H2

and

48

n = 0.268 if 5 > 1.9 H2

and f(t) is an assumed linear function as suggested earlier by Brauer and Kriegel [ 71 .

The constant K in eqn. (7) was obtained from all the previous results for velocity, hardness, particle size and concentration and with the help of the least-squares method. The final equation expressing wear for cast iron is

AW,, = 1.342 X 1O-5CV 0.682 0.616~2.39 a-cxl

-180-90 90-o!,

where n is again a function of HI/H,. The experimental and calculated values of the specimen mass loss are plotted graphically in Fig. 8, where the ideal line of unit slope is shown.

2.4

2.0

T

d IJ 1.6

z

2 2 1.2

a:

B

.8

.4

L

0 .4 .8 1.2 1.6 2.0 2.4

CALCULATED AW, GKAM

Fig. 8. Relationship between measured and calculated values of AN’.

5. Conclusions

(1) A review of previously used abrasion testing facilities produced a jet abrasion test facility that has a simple design, although each wear param- eter can be measured.

49

(2) Experimental and calculated results for cast iron as an example of hard and brittle materials are comparable on a specimen mass loss basis.

(3) From the analytical model, the prediction of wear resistance for any sample of cast iron under any running condition can be obtained.

References

1 R. G. Bayer, P. A. Engel and S. L. Sirico, Impact wear testing machine, Wear, 19 (1972) 343 - 354.

2 W. J. Head, L. D. Lineback and C. R. Manning, Modification and extension of a model for predicting the erosion of ductile materials, Wear, 23 (1973) 291 - 298.

3 J. H. Neilson and A. Gilchrist, Erosion by a steam of solid particles, Wear, 11 (1968) 111 - 122.

4 J. E. Goodwin, W. Sage and G. P. Tilly, Study of erosion by solid particles, Proc., Inst. Mech. Eng., London, 184 (Part 1) (15) (1969 - 1970) 279 - 292.

5 S. S. Nekrasov and N. E. Ofengenden, Study of the abrasion resistance of plastics, cast basalt and rubbers in hydraulic mixtures, Sov. Plast., (11) (1961) 30 - 32.

6 I. Finnie, A. Levy and D. McFadden, Fundamental mechanisms in the erosive wear of ductile metals by solid particles, Erosion: Prevention and Useful Application, ASTM Spec. Tech. Publ. 664, 1979, pp. 36 - 58.

7 H. Brauer and E. Kriegel, Untersuchungen iiber den Verschleiss von Kunststoffen und Metallen, Chem.-Zng.-Tech., 35 (10) (1963) 697 - 707.