Wear 256 (2004) 937–947

Comparison of predicted and experimental erosion estimatesin slurry ducts

R.J.K. Wooda,∗, T.F. Jonesb, J. Ganeshalingamb, N.J. Milesb

a Surface Engineering and Tribology Group, School of Engineering Sciences, University of Southampton, Highfield, Southampton SO17 1BJ, UKb School of Chemical, Environmental and Mining Engineering, The Nottingham Mining and Minerals Centre,

University of Nottingham, University Park, Nottingham NG7 2RD, UK

Received 22 April 2003; received in revised form 24 July 2003; accepted 9 September 2003

Abstract

Computational models for the impact velocity and impact angle in a bend have been successfully applied to the flow field and validatedby electrical resistance tomography (ERT) to confirm the position of particle burdens. Particle impact parameters from this work have beenused as inputs to erosion models to predict wall wastage rates in aRc/D = 1.2 bend and a upstream straight pipe section. The location ofdamage and the levels of wall wastage agree well with those obtained from erosion trials conducted on a full scale loop rig with AISI 304stainless steel pipework. Well-distributed but asymmetric particulate flows of sand in water have been studied at average flow velocities of3 m s−1. Predicted erosion damage levels and locations are compared to non-destructive ultrasonic and gravimetric measurements as wellas wall thickness measurements made with a micrometer after cutting the pipe and bends sections.© 2003 Elsevier B.V. All rights reserved.

Keywords: Erosion; Pipes; Bends; Sand slurry; Microcutting; Wall wastage; CFD; Models

1. Introduction

There is little disagreement that erosion damage in con-ventional pipelines is costly and inconvenient. In the UKalone, the Department of Trade and Industry in 2000 put ero-sion costs at approximately £20 million per annum and Jost[1] estimated wear-related figures as high as 1.5% of GNP.

The smooth interior profile of cylindrical ducts encour-ages settling of particles from low-speed slurries. Designerstend to specify a flow velocity which will entrain settlingparticles into the flow but this is usually higher than neededsatisfactorily to transport the slurry. Pumping power require-ments are high as is the propensity for components to wear.Bends act like concave mirrors and reflect the flow (with anyentrained particles) leading to localised wear.Fig. 1 illus-trates the geometry of straight and curved ducts and the po-sition of asymmetric particle burdens within the bore. Pleasenote the definitions of the pipe invert, intrados and extradosas applied to the geometry of ducts, appear in the Glossaryand in Fig. 1. It is tempting to view the lower part of theduct (pipe invert) as an obvious location for accelerated wearin straight pipes and the outer extreme (extrados) for accel-

∗ Corresponding author. Tel.:+44-1703594881; fax:+44-1703593230.E-mail address: rjw3@soton.ac.uk (R.J.K. Wood).

erated wear in bends, but this is not always the case. Thediscussion which follows sets the scene for a wider consid-eration of particle positions and wear.

1.1. Particle positions and wear

Most published results refer to maximum wear at the bot-tom of a pipe. Little information is available about the dis-tribution or circumferential location of erosion modes. Asimple model for erosive wastage at a particular orientation,θ, in the duct includes primarily a factor for the energy dis-sipation of impacting particles plus friction and directionalimpingement at that orientation[2]. There must also be amultiplicative factor for the meansupply of impacting par-ticles at that orientation (Nθ particles min−1). Factors suchas impact angle and velocity-based corrections for plasticshear deformations are left for later discussion.

Asymmetric slurry flow can be modelled as two layersfollowing Wilson [3,4]:

• a contact layer where moving particles are in continuous,indirect or sporadic contact with the duct,

• a suspended layer.

The principle of the supply of energetic particles, elabo-rated above, leads us to consider three possible modal posi-tions for wear:

0043-1648/$ – see front matter © 2003 Elsevier B.V. All rights reserved.doi:10.1016/j.wear.2003.09.002

938 R.J.K. Wood et al. / Wear 256 (2004) 937–947

Nomenclature

Ck cutting characteristic velocity (m s−1)Cv volume fractionD internal diameter of bend (m)Dk modified deformation characteristic

velocity (m s−1)Ef deformation erosion factor (J m−3)Ep Young’s modulus for the particle (N m−2)Et Young’s modulus for the target (N m−2)Er erosion rate (�m3 min−1)L pipe length (m)Li input leg of bend (m)Lo output leg of bend (m)M total mass of particlesMp particle mass (kg)ML target mass loss (kg)n velocity ratio exponentNθ particle supply at a given positionθ on

the interior surface of a duct, particles (s−1)qp Poisson’s ratio for particleqt Poisson’s ratio for targetQv volumetric flow rate of solids (m3 s−1)rp particle radius (m)R pipe radius (m)Rc centreline radius of bend (m)Rf roundness factor for particle (value 0–1)Up particle velocity (m s−1)Wc cutting wear rate (m3 per impact)Wd deformation wear rate (m3 per impact)Wt total wear rate (m3 per impact)Y yield stress of target (N m−2)

Greek lettersδ wall penetration (mm)δmean mean penetration depth (mm)∆ difference in wall thickness (mm)θ orientation of a point on the interior surface

of the bore of a duct (◦)±θi peripheral orientation of interface between

contact and suspended layers (◦)ρp density of particle (kg m−3)ρt density of target (kg m−3)σ plastic flow stress for target (N m−2)

1. θ ≈ ±θi : The boundary region between the contact andsuspended layers where particle velocities can be highand the particle supply can also be high.

2. θ ≈ 0◦: The base of the contact layer where particlesupply is relatively high but particle velocities, andtherefore erosion efficiencies, are relatively low.

3. θ ≈ 180◦: The top of the suspended layer where par-ticle velocities are relatively high, but particle supplycan be extremely low.

Fig. 1. Asymmetric suspensions in straight sections and bends, showingkey circumferential positions.

Roco and Cader[2] show wear modes at the boundaryregion when velocities are near or below the deposition ve-locity, and at the base of the contact layer when velocitiesequal or exceed the deposition limit. At higher slurry veloc-ities, wear is distributed around the periphery of the bore, inconcert with the asymmetric distribution of particle supply.In a loop operating just above the deposition velocity, thelocal particle velocity and particle supply will vary aboutthe mean values. A mixture of wear modes is therefore tobe expected.

1.2. Slurry flow patterns

The pattern of slurry flow is important in determiningthe kinetic energy and supply of particles at a point on theperiphery of a pipe. Slurry flow patterns are commonlycategorised into several identifiable states of suspension orpartial settling. At one extreme is the stationary bed andat the other is symmetrically suspended flow. Govier andAziz [5] subdivided the transition into five main categorieswhile Chien[6] classified sand/water mixtures into an evengreater number of categories.

Williams and Beck[7] show that electrical resistance to-mography (ERT) can be used to determine these patterns ex-perimentally. Evenly spaced electrodes around the pipe andin contact with the slurry are used to construct a grid of con-ductivity measurements which can be used to plot contoursof equal slurry concentration. It is important to note that thetechnique does not measure the density of each particle butassigns a concentration to regions of the cross-section. Forstudying the patterns of flow this presentation is ideal.

Flow pattern determinations were performed at the Uni-versity of Nottingham using a straight duct and white beadsof median size 1.75 mm (±0.257 at 1 s) and median relativedensity of 1.45 (±0.005). Measurements were made using acommercial ITS-P2000 ERT system. A current injection pro-tocol based on excitation of adjacent electrodes using 10 kHzalternating current was used. The gains of pre-amplifiers forall measurement were acquired by automatic calibrations.Images were reconstructed using a sensitivity-coefficientback-projection algorithm after Wang et al.[8]. The tomo-graphic concentration curves were obtained from a ring of16 electrodes which were evenly spaced around the periph-ery of the pipe at the measurement plane. Concentrationfractions were obtained from conductivity measurements be-tween electrodes.Fig. 2 shows the concentration contoursobtained at three distinct flow patterns: astationary bed in

R.J.K. Wood et al. / Wear 256 (2004) 937–947 939

(a) 0.5 m/s (b) 1 m/s (c) 2 m/s

Fig. 2. ERT data for the flow of beads in a straight duct (2 mm beads, relative density 1.45, nominal in situ concentration 5% (v/v)): (a) stationary bedpattern; (b) moving bed pattern; (c) asymmetric suspension.

which there was no discernable movement of packed beadsat the pipe invert, amoving bed in which settled beads wereprogressing along the pipe and lastly an asymmetric suspen-sion. The diagrams inFig. 2 demonstrate the possible wearmodes mentioned above. The use of 16 electrodes caused thegrid to be fairly coarse. The slight misorientation of thesediagrams was found to be caused by digitisation effects.

1.3. Wear from flow of slurry

Slurry erosion is complex and good models have yet to bedeveloped. This is certainly true for erosion of pipelines andis compounded by the fact that most published literature onthe wear of pipework relates to pneumatic conveying sys-tems. Meng and Ludema[9] quote 33 independent parame-ters in a recent review of 22 erosion models and predictiveequations found in the literature. Thus, the complexity ofslurry erosion is due to the number of important variablespresent, some of these are listed inTable 1.

Table 1Some of the main variables which influence erosion

Slurry variables Component variables

Liquid: viscosity, density, surfaceactivity, lubricity, corrosivity,temperature

Bulk properties: ductility orbrittleness, hardness andtoughness, melting point,microstructure, shape androughness

Particles: brittleness, size, density,relative velocity, shape, relativehardness, concentration,particle/particle interactions

Surface properties: workhardening, corrosion layers,surface treatments, coating type,coating bond, microstructure

Flow field: angle of impingement,particle impact efficiency,boundary layer, particle rebound,degradation, particle drop-out,turbulence intensity

Service variables: contactingmaterials, pressure, velocity,temperature, surface finish,lubrication, corrosion, hydraulicdesign, intermittent slurry flows

For pipe bend erosion the problem is to develop a modeland in parallel develop a better understanding of the flowfield and particle trajectories to generate accurate inputs intothe model, and then compare the model predictions withexperimental results.

Discussion (recently in an excellent review by Kaushaland Tomita[10]) has focussed on the distribution of erod-ing particles in a cross-section through a straight duct, butunderstanding the particle trajectories throughbends is ofgreat importance if wear rates are to be predicted at specificlocations. For bends in a pneumatic conveying line, Millsand Mason[11] found significantly greater wear depth for70�m sand compared to 230�m sand. Blanchard et al.[12]attempted a two-dimensional model of particle trajectoriesin liquid flows within bends but with limited success due tothe inability of the model to predict secondary flows. At theUniversity of Southampton, Forder[13] has studied solidparticle trajectories in liquid flows in pipe bend geometrybefore modelling more complex flows in choke valves us-ing a commercial computational fluid dynamics (CFD) code.Forder[13] extended this commercial CFD programme toinclude a predictive erosion element based on a combineddeformation and cutting erosion model and this will be dis-cussed in detail in the following section. The model wastested by comparing the predicted wear rates and wear lo-cations in pipe bends with the experimental air/sand erosionresults of Bourgoyne[14], with excellent correlations. Fur-ther experimental wear results for pipe bends could be usedto test this approach in the future such as reported by Kinget al. [15] and Wiedenroth[16] for a limited range of bendgeometries, materials and liquid–solid flows. The redistri-bution of solids inside horizontal bends for multisized par-ticulate slurries has been experimentally studied by Ahmedet al.[17]. Conclusions from this work point to the redistri-bution of large particles outwards to be the likely cause ofrapid bend erosion.

940 R.J.K. Wood et al. / Wear 256 (2004) 937–947

2. Component erosion models

The prediction of erosion rates and locations is difficultfor equipment handling particle-laden process streams. Theevolution of erosion models and commonality between themhas been reviewed by Meng and Ludema[9], Wood et al.[18] and Hutchings[19].

Recent erosion models recognise the two erosion mech-anisms of cutting and deformation erosion, with discretemodels representing each, e.g. Forder[13]. A cutting ero-sion model was first proposed by Finnie[20] and latermodified by Hashish[21]. For low impact angles (typi-cally 0–40◦) of relatively hard particles on ductile targets,cutting is likely when the shear stresses induced by theimpact exceed the shear strength of the target. A defor-mation model, Bitter[22], is thought applicable at higherimpact angles (30–90◦). A stress field is generated withinthe contact inducing plastically deformed sub-layers wherethe stress (typically sub-surface) exceeds the yield strengthof the target material. This leads to erosion by delamina-tion, micro-cracking and coalescence of voids to producesurface fragmentation. These models operate in parallel andthe summation of the two presents the overall volumetricerosion per impact and can be given by:

Wt ={

100

2√

29r3p

(Up

Ck

)n

sin 2α√

sinα

}

+{

Mp(Up sinα − Dk)2

2Ef

}(1)

where

Ck =√

3σR0.6f

ρp(2)

and

Dk = π2

2√

10(1.59Y)2.5

(Rf

ρt

)0.5[

1 − q2p

Ep+ 1 − q2

t

Et

]2

(3)

andn = 2.54 for carbon steel, Hashish[21], Ef = 1.9×1010

for AISI 4130 steel.This model, as used by Forder[13], includes important

variables such as particle shape (roundnessRf ) and materialproperties for both particle and target that are typically notconsidered by earlier simple models.

This approach is similar to that of a recent paper by Li andShen[23] which describes a computer simulation of pipebend erosion in pneumatic systems conveying low concen-trations of granular material. Forder[13] and Li and Shen[23] use computer modelling to determine the velocity fieldof the particles for given boundary geometries, particle sizeand density, particle loading, fluid type and driving pressuregradient. This information is placed into an erosion modelbased onEq. (1). The empirical constants in these mod-els were set by[13,23] to agree with experimental work by

Table 2Values of parameters used inEqs. (1)–(3)

Parameters Value

Volume flow rate,Qv (m3 s−1) 0.01Pipe bore (m) 0.078Particle density,ρp (kg m−3) 2670Roundness factor,Rf 0.5Plastic flow stress,σ (Pa) 1.00E+09Yield stress of target,Y (Pa) 3.20E+08Target density,ρt (kg m−3) 7850Target Poisson’s ratio,qt 0.3Particle Poisson’s ratio,qp 0.23Particle Young’s modulus,Ep (Pa) 5.90E+10Target Young’s modulus,Et (Pa) 2.07E+11Deformation erosion factor,Ef (J m−3) 1.9E+10Particle radius,rp (�m) 500Velocity ratio exponent,n 2.54Particle velocity on entry,Up (m s−1) 3.0Particle mass,Mp (kg) 1.4E−06

Shimizu et al.[24] using copper and glass or established us-ing a slurry jet impingement rig at various jet angles up to ajet velocity of 35 m s−1 on stainless steels and sintered tung-sten carbides, Forder[13]. Wallace et al.[25] used a similarapproach coupling a CFD package with an alternative ero-sion model, with fewer variables, based on that of Neilsonand Gilchrist[26]. These models still need to be verifiedby experiment and in the case of the Neilson and Gilchristmodel at least four sets of experimental data are requiredobtained at different impingement angles and velocities toevaluate the empirical constants.

2.1. Computational simulations

Computational simulations have been made of a straightpipe and bend that evaluate the impact angle and velocitydistribution over the internal surfaces. These values are in-corporated into the Bitter–Hashish component erosion mod-els based onEq. (1)to predict erosion patterns within thesecomponents. FLUENT v5.4 CFD code1 with an embed-ded multi-phase (algebraic slip) model was used. A bendof 0.078 m bore diameter and aRc/D ratio of 1.2 with a0.078 m bore diameter straight pipe section 0.56 m in lengthlocated directly upstream of the bend was modelled. Bothcomponents were modelled in a horizontal plane. The liq-uid phase was set to typical values for water (viscosity=0.0009 Pa s, density= 1000 kg m−3), the solids had a den-sity of 2670 kg m−3 and a size of 1 mm. Both liquid and solidentry velocities were set to 3 m s−1 with a uniform solid dis-tribution. A standardk–ε turbulence model was used withstandard wall functions and zero roughness. Impact angles(α) were defined as those from the tangent to the bend wall.Table 2summarises the parameters used.

1 Fluent Inc., Centerra Research Park, 10 Cavendish Court, Lebanon,NH 03766, USA.

R.J.K. Wood et al. / Wear 256 (2004) 937–947 941

180˚

90˚

0˚

270˚

-16-14

-12

-10-8

-6

-4-2

0

24

6

0 90 180 270 360

in plane angle ( )

Imp

act

ang

le (

˚)

0 plane

45 plane

pipe cross-section

0

2

4

6

8

10

12

0 90 180 270 360

in plane angle (˚)

San

d v

olu

me

(%)

0 plane

45 plane

0

0.5

1

1.5

2

2.5

3

0 90 180 270 360

in plane angle (˚)

Imp

act

velo

city

(m/s

)

0 plane

45 plane

0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

5

0 90 180 270 360

in plane angle (˚)

Ero

sio

n r

ate

(mm

3 /m

inu

te) 0 plane

45 plane

Fig. 3. Computational results.

The wear rates can be calculated by using the modifiedHashish–Bitter model,Eq. (1). The variation of the overallwear rateWt with impingement angleα is plotted elsewhere,Wood et al.[18], and predicts maximum erosion to occur atslightly different angles to that of Neilson and Gilchrist.

The results from the coupled CFD and erosion model areillustrated inFig. 3. Particle impact angles and velocities ata plane midway along the straight section and perpendicu-lar to its axis (labelled 0◦ plane within they–z plane) wereevaluated as well as for a plane at 45◦ or midway aroundthe bend. The impact angles, sand volume, impact velocityand erosion rates are plotted for both the 0◦ plane within thestraight section and the 45◦ plane within the bend. Although

AlternativeRoute

MagneticFlow MeterCentrifugal

Slurry Pump

Straight 43

Bend 65

StirredTank ~3m3

2.64m

2.48m

4.68m

Fig. 4. Details of the experimental loop.

secondary recirculating flows (negative impingement angles)are evident between the in-plane angles of 90◦ and 270◦ theparticle loading and impact velocities between these anglesare such that little wear is predicted with most damage oc-curring 270◦ and 90◦.

3. Experimental loop

The experimental loop comprised horizontal straights andbends,Fig. 4. The rig could be broken down into sectionsto allow inspection after each test. For each test, sand–watermixtures at 10 vol.% were pumped through the rig at a flow

942 R.J.K. Wood et al. / Wear 256 (2004) 937–947

C

AG

E

00.10.20.30.40.5

0˚ 90˚ 180˚ 270˚

all straight pipe sections

0˚ 90˚ 180˚ 270˚

DCBA

G

E

00.10.20.30.4

0˚ 90˚ 180˚ 225˚ 270˚ 315˚

Circumferential position of wear mode:Circumferential position of wear mode:all bends

Pro

bab

ility

Pro

bab

ility

0˚ 90˚ 180˚ 225˚ 270˚ 315˚

Flow into Page

A

B

C

DE

F

G

H

0˚

90˚

180˚

270˚

(c) (d)

Wear mode (0.76 µm/h)

(a) (b)

Fig. 5. Wear modes and their distribution over straight and curved ducts: (a) measurement points (not all were used on every plane); (b) measurements inone plane (ACEG) on a straight duct showing wear mode at position C; (c) distribution over all straight pipe sections (13 planes in total); (d) distributionover all curved ducts (18 planes in total).

velocity of 3 m s−1 over 30 min periods, with diversion toan alternate loop of the same geometry for 30 min. The testswere repeated to a total of 210 h with mean∼1000�m di-ameter sub-rounded to sub-angular silica sand (D14/25 sizerange 500–1400�m and hardness 1100 Hv) from HepworthMinerals and Chemicals Ltd., Leighton Buzzard, UK. Thetwo test components, straight 43 and bend 65, are shownin Fig. 4 because wear patterns in these components willbe described below. These patterns will be compared to thepredicted wear rates from the CFD and erosion model basedon these component geometries.

3.1. Wastage measurements

Following each test, pipe wall losses were measured bya combination of techniques. The rig was dismantled sothat each section could be weighed, measured by ultrasonicthickness probe and visually inspected by endoscope. Beforethis could be done, the sections were cleaned with warm5% (v/v) citric acid to remove calcite deposits on internalsurfaces. This was necessary to allow the change in mass ofthe pipe sections to be attributed solely to wear. The pipeswere then allowed to dry in air prior to weighing.

• Gravimetric losses in each pipe were obtained by sub-tracting its weight before and after each test. Each pipesection was placed on an aluminium frame which dis-tributed its weight between four electronic balances. Eachbalance was nominally capable of weighing up to 8.1 kgto the nearest 0.01 g. The outputs of the balances werecontrolled by means of an RS232 interface to a personalcomputer.

• Ultrasonic thickness gauges (Krautkramer type DM4 DL)were used to measure the thickness at a selection of eightcircumferential points labelled ABCDEFGH in a clock-wise direction. Point A was at the top of the pipe (or180◦ in-plane angle), point C was at the 3 o’clock posi-

tion and so on. The planes at which the surveys were car-ried out were defined so that results could be comparedfrom test to test. Straight sections were tested at the car-dinal points ACEG. Bends were tested at ABCDE andG to capture the outside of the bends, the top and bot-tom point and one point (G) at the inside of the bends.Fig. 5(a) shows the measurement points andFig. 5(b)shows a typical plane through a pipe section showingthe “mode” or maximum of the wear distribution at thatplane.Fig. 5(c) and (d) show the distribution of the wearmodes (maximal wear points) in straight sections andbends.

Following the series of tests, some components were de-structively tested and measured by ball ended micrometerat the locations at which ultrasonic measurements had beentaken. The results of these tests on the bend are shown inFig. 6. Ultrasonic readings showed a small but significantbias. This was explained as follows. The contact area, or“zone” of the ultrasonic beam was different from the area ofcontact of the ball-ended micrometer. Both instruments aremost accurate when estimating the thickness of a rectilin-ear piece. When the increased contact zone of the ultrasonicbeam is superimposed onto the curved surface of the pipe,a slight over-estimate of the thickness of the pipe occurs.The smaller contact area of the micrometer caused a smallerover-estimate. When the ultrasonic readings were comparedwith the micrometer readings the former had a positive bias,i.e. the ultrasonic estimates were higher than the micrometerestimates. The calibration of the ultrasonic instrument wasdone on a square flat test piece of the same steel as that usedfor the pipes and did not take pipe curvature into account,so the full regression line inFig. 6 incorporates a smallintercept. The ultrasonic transducer was therefore expectedto read slightly high for lower wastage in straight pipe 43and slightly low for high wastage in bend 65. The modalbias inFig. 6(c) differed slightly from the regression value.

R.J.K. Wood et al. / Wear 256 (2004) 937–947 943

4.0

4.5

5.0

5.5

6.0

4.0 4.5 5.0 5.5 6.0

Micrometer (mm)

Ultr

ason

ic(m

m)

0.0

1.0

2.0

3.0

4.0

5.0

6.0

0.0 1.0 2.0 3.0 4.0 5.0 6.0

Micrometer (mm)

Ultr

ason

ic (

mm

) Data

y = 0.9731x

y = 1.0064x - 0.177

0

1

2

3

4

5

6

70

0.04

0.08

0.12

0.16 0.2

0.24

0.28

0.32

0.36 0.4

Micrometer-Ultrasonic Reading

Fre

qu

ency

(a) (b)

(c)

Fig. 6. Comparison between ultrasonic gauge and micrometer at the same measurement points over every measurement plane on four pipe components: (a)calibration lines with and without intercept showing crossover; (b) good correspondence at small scale; (c) distribution of deviations between micrometerand ultrasonic readings.

The standard error was±0.04 mm, giving an error range of0.09–0.17 mm (using a modal bias of 0.13 mm). Typical er-rors from micrometer readings, obtained by repeating themeasurements, were found to be in the order of±0.03 mm.The uncertainty from gravimetric readings, also obtainedby repeating measurements, was estimated at±0.09 g at1 S.E.

Destructive testing of the components also allowed a de-tailed analysis of the erosion patterns using a scanning elec-tron microscope.Fig. 7presents micrographs of the erosionpatterns from the test loop at the top (point A or 180◦) andside (point D or 315◦) at the 45◦ plane of the bend and bot-tom (point E or 360◦) of the straight pipe at the 0◦ the plane.FromFig. 5(d) it is clear from the ultrasonic measurementsthat point E (bottom) suffers the most damage while point D(side) suffers approximately half as much and point A (top)the least damage. This agrees well with erosion damage in-formation from the micrographs of each point as explainedbelow. Fig. 7(a) shows point A has a complex surface to-pography matching that of the as-received internal pipe sur-face with evidence of some erosion damage to protrudingareas.Fig. 7(b) shows point D that is within the boundary

region between the suspended and contact layers and showsa low density of microcutting scars (mainly aligned withthe flow direction). This is consistent with the particle posi-tion model described inSection 1.1. Relatively large num-bers of particles of relatively high kinetic energy impinge onthe pipe surface. In the present study, these particles haveeroded away most of the original surface features, althoughsome original features are seen in the centre of the micro-graph.Fig. 7(c) shows typical severe damage patterns foundat point E with multiple microcutting impact sites (mainlyaligned with the flow direction) and complete removal of allas-received surface features. Further details can be found inWood and Jones[27].

The comparison of gravimetric data for bend 65 andstraight pipe 43 should be done on the basis of the areaof the pipe interior which is in contact with the particleburden. The first step is to calculate the entire area of thepipe interior or “wetted” area. For the purpose of discus-sion we assume that the area in contact with the particleburden is 25% of the wetted area. It is not absolutely nec-essary to make such an assumption for comparison of twogravimetric estimates, but does yield more realistic data for

944 R.J.K. Wood et al. / Wear 256 (2004) 937–947

Fig. 7. SEM micrograph showing wear pattern damage of the bend after 210 h at: (a) point A or 180◦; (b) point D or 315◦; (c) the straight pipe at pointE or 0◦.

comparison with other methods. For both sections,

δmean= ML

ρt × 0.25× wetted area= 4ML

ρt × wetted area(4)

The wetted areas can be obtained fromEqs. (5) and (6):

wetted area= 2πRL (for straight 43) (5)

whereR is the pipe radius andL the pipe length

wetted area= π2

4

{R2

c −(

D

2

)2}

+ 2πR(Li + Lo)

(for bend 65) (6)

whereRc is the bend radius, andLi , Lo the input and outputlegs.

Table 3shows gravimetric data from the two pipe sectionsconverted to a mean penetration as described above.

The micrometer results for both test geometries are pre-sented inFig. 8. A total of 57 measurements were taken foreach plane (0◦ for the straight and 45◦ for the bend). To al-low for variations in wall thickness induced in componentmanufacture values obtained at planes of no or little erosion

(just downstream of inlet weld for both components) havebeen subtracted from the in-plane values to generate a differ-ence in wall thickness (∆). Not all the wall variations in themanufacture of the bend can be removed by this method asseen inFig. 7(a) with positive differences in wall thickness.

Fig. 8(a) indicates that, for the bend, an asymmetricerosion damage pattern has occurred between the in-planeangles of 225◦ and 90◦ (points BCDEFG). Although, asexpected, it is noted that there is more (approximately 1.5times) erosion of the outer bend radius between in-planeangles 225◦ and 360◦ (points BCDE) compared to the in-ner radius erosion that occurs between 0◦ and 90◦ (pointsEFG). The CFD code predicts a fairly symmetrical erosionpattern, seeFig. 3.

Table 3Gravimetric measurements,ML , after 210 h

Straight 43 Bend 65

Gravimetric (g) 4.75 9.45Wetted area (m2) 0.14 0.0613Mean penetration over 25%

wetted area,δmean (mm) fromEq. (6)

0.017 0.077

R.J.K. Wood et al. / Wear 256 (2004) 937–947 945

-1.50

-1.00

-0.50

0.00

0.50

1.00

0 90 180 270 360

In plane angle (˚)

∆ w

all t

hic

knes

s (m

m)

-0.25

-0.20

-0.15

-0.10

-0.05

0.00

0.05

0.10

0.15

0.20

0.25

0 90 180 270 360

In plane angle (˚)

∆ w

all t

hic

knes

s (m

m)

(a)

(b)

Fig. 8. Difference in micrometer derived wall thickness between unerodedand eroded areas: (a)∆ wall thickness measurements for the bend sectionat plane 45◦; (b) for the straight pipe section at plane 0◦.

A less clear trend in measured penetration depth is seen forthe straight pipe,Fig. 8(b). Although, erosion has occurredbetween 290–360◦ and 30–80◦ that is similar to the erosionlocations predicted inFig. 3. The maximum and mean valuesof these results are compared with the other measurementsand CFD predictions inTable 4.

A histogram of the results is given inFig. 9. The ratiosof the mean values of bend to straight erosion are listedin Table 5. It should be noted that whilst comparisons be-tween the ultrasonic, micrometer and CFD predictions are

Table 4Summary of measured, calculated and predicted erosion penetration depths

δ measurements Bend 65 Straight 43

Ultrasonic thickness (mm)Maximum 0.41 0.40Mean 0.11 0.09

Micrometer thickness (mm)Maximum 0.70a 0.19Mean 0.21 0.02

CFD (mm)Maximum 0.72 0.113Mean 0.24 0.035

Gravimetric (mm)Mean 0.077 0.017

a Corrected for manufacturing processes that thin the bottom andthicken the top walls.

Fig. 9. Comparison of wastage estimates (using results fromFig. 3and thewetted areas fromEqs. (5) and (6)) with micrometer and ultrasonic mea-surements from pipe loop experiments at 3.0 m s−1, 1 mm sand, 10 vol.%sand concentration after 210 h.

Table 5Ratios of bend to straight mean erosion

Mean bend/mean straighterosion ratio

Ultrasonic thickness 1.2Micrometer thickness 10.5CFD 6.9Gravimetric (all planes) 4.5

in-plane at one section, the gravimetric results reflect anaverage loss of all planes and are unable to detect ero-sion maxima.Fig. 10 plots the measured (ultrasonic andmicrometer) results against the CFD predicted results andshows good agreement between the micrometer and CFDdata while the ultrasonic measurements show some scatter.The ultrasonic probe appears to yield higher values of wallthinning on the straight compared to those predicted whilethe reverse is true for the bend. Thus, caution should be

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0 0.2 0.4 0.6 0.8

CFD predicted erosion penetration (mm)

mea

sure

d e

rosi

on

pen

etra

tio

n (

mm

)

UltrasonicMicrometer

Fig. 10. Comparison between CFD predicted results (using results fromFig. 3 and the wetted areas fromEqs. (5) and (6)) with micrometer andultrasonic measurements from pipe loop experiments at 3.0 m s−1, 1 mmsand, 10 vol.% sand concentration after 210 h.

946 R.J.K. Wood et al. / Wear 256 (2004) 937–947

used when interpreting ultrasonic thickness data from pipesystems.

4. Conclusions

• Sand–water experiments using a 78 mm bore pipe loopsystem to investigate damage to a straight pipe and bendfound that wear maxima were distributed to either sideof the pipe bore for the straight pipe, in contradiction ofthe common assumption of predominant wear at the baseof a pipe. In the bend there was an expected increase inwear at the outsides of the bends compared to the in-sides, but also significant wear occurred at the base of thebend.

• The predominant erosion mechanism of the straight pipeand bend surfaces was microcutting from low angle solidparticle impingements. Lower densities of cutting scarswere seen at points within the boundary region betweenthe suspended and contact layers of sand while higherdensities were evident at points within the base of thecontact layer.

• Good agreement was found between the in-plane wallpenetration depths obtained using a ball ended micrometerand the coupled CFD and Hashish–Bitter erosion modelfor both straight pipe and bend components.

• Comparison between the coupled CFD/erosion modelsand micrometer/ultrasonic measurements of the straightpipe and bend show reasonable agreement between pre-dicted and experimental erosion patterns.

• Interpretation of ultrasonic thickness data from pipe sys-tems was hampered by higher scatter in the data comparedto the micrometer data.

• Gravimetric measurements, corrected for wetted area,show the bend to erode at 4.5 times that of the upstreamstraight pipe but gravimetric measurements of this kindare not meaningful in the context of life prediction as theyrefer to a ‘smeared-out’ estimate of erosive wear. Themaximum erosion rates and wear locations are importantfor pipe lifetime and failure prediction.

Acknowledgements

The authors would like to thank Mr. Tony Gospel of theUniversity of Nottingham and Dr. Julian Wharton of theUniversity of Southampton for their help throughout thisproject.

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Glossary

Wear distribution: The pattern of wear on the circumference of a ductcross-section, viewed as a distribution.

Wear mode: The position on the circumference of pipe cross-sectionswhere the incidence of wear damage is greatest.

Pipe invert: The pipe invert is the region at the bottom of the bore wheresettling particles tend to gather.

Extrados: The extrados is the outermost extremity of a curved duct.Intrados: The intrados is the innermost extremity.