\

CO NBS SPECIAL PUBLICATION 394

U.S. DEPARTMENT OF COMMERCE / National Bureau of Standards

The Role of Cavitation

in

_ Mechanical Fdilur

MFPG
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0 1974

0UUQ-.

MFPG

The Role of Cavitation in Mechanical Failures

Proceedings of the 19th Meeting of the

Mechanical Failures Prevention Group,

held at the

National Bureau of Standards,

October 31, November 1 and 2, 1973,

Boulder, Colorado 80302

Edited by

T. R. Shives and W. A. Willard

Metallurgy Division

Institute for Materials Research

U.S, National Bureau of Standards

Washington, D.C. 20234

The 19th meeting of MFPG and these proceedings were sponsored

by the Office of Naval Research, the National Aeronautics and

Space Administration, and the Institute for Materials Research

of the National Bureau of Standards.

U.S. DEPARTMENT OF COMMERCE, Frederick B. Dent, Secretary

NATIONAL BUREAU OF STANDARDS, Richard W. Roberts, Director

Issued April 1974
Library of Congress Cataloging in Publication Data

Mechanical Failures Prevention Group.

The role of cavitation in mechanical failures.

(National Bureau of Standards Special Publication,

39*0

Papers presented at the meeting, held Oct. 31-Nov. 2,

1973 in Boulder, Colo.

Supt.'Of Docs. No.: C 13. 10.39^

1. Erosion of metals —Congresses. 2. Cavitation

Congresses. I. Shives, T. R. , ed. II. Willard,

William A. , ed. III. Title. IV. Series: United

States. National Bureau of Standards. Special Publication, 39k.

QC100.U57 no. 39^ [TAU60] 389'. 08s [620.1'6'6] 7^-7057

National Bureau of Standards Special Publication 394

Nat. Bur. Stand. (U.S.), Spec. Publ. 394, 183 pages (Apr. 1974) CODEN: XNBSAV

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 1974 For sale by the Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402 (Order by SD Catalog No. C13.10:394). Price $2.25
FOREWORD

The 19th meeting of the Mechanical Failures Prevention Group was held

October 31 and November 1 and 2, 1973 , at the National Bureau of

Standards in Boulder, Colorado. Under the chairmanship of Professor

Elmer E. Klaus, the MFPG technical committee, Mechanisms of Failure,

organized the program for the meeting. We commend Professor Klaus,

the session chairmen, and especially the speakers for an excellent

program.

In addition to the abstracts of the talks, which are presented as submitted except for very minor editing in a few cases, these "Proceedings" include the discussions of the talks. These discussions

were recorded at the meeting and have been edited to improve readability.

Special appreciation should be extended to Mr. T. R. Shives and Mr.

W. A. Willard of the NBS Metallurgy Division for their editing, organization, and preparation of the "Proceedings", Mr. Kenneth Armstrong

of the Boulder Special Activities Office of NBS for the meeting arrangements, Mr. H. C. Burnett of the NBS Metallurgy Division for general

coordination, and the entire staff of the Metallurgy Division and

the Institute for Materials Research for their assistance in many

ways. Special thanks are accorded Miss Bronny Evans of the Metallurgy

Division for her diligent efforts in transcribing the recorded discussions.

ELIO PASSAGLIA

Executive Secretary, MFPG

Chief, Metallurgy Division

National Bureau of Standards

III
TABLE OF CONTENTS

FOREWORD

SESSION I

lo Physics Associated with Cavitation Induced Material

Damage. F D B. Peterson

2o Role of Physical Properties of Liquids in Cavitation

Erosion. A. Thiruvengadam

3. The Role of Mechanical Properties in Cavitation

Erosion Resistance. G. C. Gould

h. Recent Theories of Cavitation Damage Including NonSymmetrical Bubble Collapse Effects. F. G. Rammitt

SESSION II

1. Some Practical Examples of Cavitation Erosion and

Their Prevention „ A. F. Conn

2. Examples of Oil Cavitation Erosion in Positive

Displacement Pumps. J. A„ Halat* and G. 0. Ellis*

3. Microscopic Investigation of Cavitation Erosion Damage

in Metals. J. V. Hackworth* and W. F. Adler

k. Lubricant Pressure Rippling in Dynamic Hertzian Contacts Induced by Surface Roughness. T. E„ Tallian

and J. I. McCool*

SESSION III

1. Asperity Lubrication and Cavitation in Face Seals.

J. A. Walowit

2. Effect of Cavitation on Fluid Stability in Polymer

Thickened Fluids and Lubricants. J. L. Duda and

E. E. Klaus*

3. Effects of Viscoelasticity on Cavitation in Drag

Reducing Fluids., R. Y. Ting

h. Progress and Problems in Erosion Prediction,,

F 0 J. Heymann

IV
SESSION rv

1. Potential of Thin, Sputtered Films as Erosion Resistant Protective Coatings.

Part 1. Sputter -Coating and Film Characteristics.

K. Gentner

Part 2. Erosion and the Gas Turbine Engine.

J. E. Newhart

2„ Rain Droplet Erosion Mechanisms in Transparent Plastic

Materials. G. F. Schmitt, Jr.

3. Analysis of Navy Radome Failure Problems. G. J„ Tatnall

and K. Foulke*

k. Government -Industry Data Exchange Program (GTDEP).

E. T. Richards

* Indicates speaker when a paper had more than one author.

LIST OF REGISTRANTS AT THE 19th MFPG MEETING

V
ABSTRACT

These proceedings consist of a group of sixteen submitted papers and

discussions from the 19th meeting of the Mechanical Failures Prevention

Group which was held at the National Bureau of Standards on October 31

and November 1 and 2, 1973. The central theme of the proceedings is the

role of cavitation in mechanical failures.

Key words: Bubble collapse; cavitation; cavitation damage; cavitation

erosion prevention; erosion; surface roughness.

VI
SESSION I

!

Chairman: E. E. Klaus

The Pennsylvania State University

PHYSICS ASSOCIATED WITH CAVITATION INDUCED MATERIAL DAMAGE

F. B. Peterson

Naval Ship Research & Development Center

Bethesda, Maryland 2001*4-

The role of cavitation in mechanical failure is extremely complex

and in many aspects is very imperfectly understood. Some of the most common types of material damage associated with the presence of cavitation are surface material removal, delamination and structural vibration.

This occurs in external flows such as on propellers, hydrofoils, and

high speed non-lifting surfaces. In internal flows, pipe bends, inlets,

constrictions, pumps and turbines are typical. Nominally nonflowing

liquids are also susceptible in, for example, strong acoustic fields and

high energy particle detectors.

For flowing systems, Bernoulli's equation shows how a local pressure

is reduced as the fluid's velocity is increased. At sufficiently high

velocities, a tension can actually develop and this has, in fact, been

demonstrated experimentally. Once the pressure is reduced below the

fluid vapor pressure a vapor cavity can be nucleated. Various aspects

of this process are simply shown by considering the flow over a lifting

surface. When the flow is completely unseparated, cavities will be

generated in the low pressure region and will translate along the surface. These cavities are frequently quite large and can produce

material damage. If the flow is separated then a stationary cavity is

produced which remains fixed on the surface. When this fixed cavity

becomes sufficiently large the downstream portion breaks up into many

small vapor cavities and these nominally spherical cavities then collapse

in the high pressure region. It is this process that typically can

produce very intense erosion, noise and vibration. It should be noted,

of course, that this separated flow can be the result of both gross

boundary layer separation and isolated local roughness elements on the

3
body. When the cavity length is increased to the point where the collapse occurs downstream of the body, then damage will typically occur only

on a downstream surface that is in the collapse region. The basic ideas

just discussed can be applied to both internal and external flows with

only minor modifications for most situations.

Cavitation in acoustic fields is to some extent similar to the unseparated flow situation just discussed because both require a continuous

supply of "weak spots", or nuclei, to produce the cavities. Herein lies one of the key aspects with regard to the extent of material damage for

various fluids in cavitating systems. If separated flow exists, one nucleus may be all that is required to produce a fixed cavity and the

subsequent generation of many smaller vapor cavities. Once this fixed

cavity is formed, the local flow properties of the system are changed

and the cavity can be eliminated only by a significant increase in system pressure or a decrease in flow velocity (i.e., a hysteresis effect).

Conversely, when many nuclei are present, and the flow is unseparated,

then very little hysteresis is experienced.

There are many methods by which cavities in a fluid can be nucleated.

Some of the most common would include, noncondensable gas bubbles in the

liquid, noncondensable gas trapped in particulate matter or in cracks on

a solid surface, a solid with low surface-free energy, and high energy

particles. If microbubbles are present, then their path in both an acoustic field or a flowing system will be influenced by the local pressure gradients. In a flowing liquid, this is the so called "screening"

effect. If a microbubble is subjected to fluctuating pressures due to

turbulence or an acoustic field, then a dramatic increase in bubble

volume is possible. This is possible since at low pressure the surface

area is greater than during the high pressure phase and thus there is a net diffusion of gas into the bubble (rectified diffusion). In addition, nonuniform pressure fields develop microscopic flow around the

bubble and change the diffusion potential across the bubble wall (microstreaming). Thus, it is seen that the position and size of a bubble
can be highly dependent on the time and pressure history of the liquid.

When these microbubbles enter a region of local pressure equal to or less

than the liquid vapor pressure, their size will determine their stability

and thus the nucleation process.

Of particular importance in water is the role of hydrophobic particles that contain some noncondensable gas. It has been shown by several

investigators that this particulate can be produced de novo in the water

and will be stabilized and of essentially neutral buoyancy. (Summarized

in ref. 1).

In the special case of water, nucleation should be relatively easy

unless extreme purification measures are taken.

In both hydrodynamically and acoustically generated cavitation it

has been experimentally verified that cavities can be nucleated on the

solid surface bounding the fluid. This process can be initiated by,

for example, noncondensable gas trapped in a crack, gas generated from

electrochemical activity, external addition of heat, regions of low

surface-free energy, or the previous collapse of a cavity on the surface

that left a small quantity of gas or particulate matter.

Certainly with all of these various mechanisms available to nucleate

a vapor cavity it is not difficult to understand the relative ease with

which most common liquids cavitate when subjected to at most several

atmospheres of tension. Once a vapor cavity has formed its subsequent

collapse can be highly dynamic and the possibility of material damage is

present. The extent of material damage is known to be basically the

result of both the mechanical and electrochemical properties of the

solid. ^Although the details of the processes are very poorly understood, it has been shown that at the termination of the cavity collapse,

a shock wave in the liquid can be produced. In addition, when other

surfaces, or pressure gradients are present, early in the collapse an asymmetry in the initially spherical shape occurs. This leads to the

involution of one part of the surface and the formation of a liquid jet

that passes through the collapsing cavity. For sufficient symmetry of

5
shape to be maintained such that a shock is produced, the cavity must be

away from a solid surface. Thus significant attenuation of the shock

occurs before it impacts on the surface. When the cavity is initially

near or on the solid then the jet most probably represents the significant mechanism for mechanical material damage. It has been observed

that the most extensive damage is apparent when the jet strikes the solid

directly and does not have to pass through any intervening liquid.

For cavitation in an acoustic field there is now strong evidence suggesting that most damage is due to the repetitive growth and collapse of a cavity at the same location on the solid surface. This discovery may

help account for the fact that acoustically generated cavitation produces

an accelerated damage rate compared with the hydrodynamically generated

cavitation. It should also be mentioned, in relation to the previous

comments, that there has been no observed cut-off in bubble size for the

damage production.

Clearly, the actual mechanism by which a material fails, following

an impact, is neither simple nor identical for all solids. However, in

order to help obtain a better insight of the mechanical damage process, even simplified views have their merit - particularly if no more exact

explanation is available. As a first approximation, solids can be con- sidered to fail first either in shear (ductile material) or in tension

(brittle material). For materials that nominally remain ductile it can

be shown that there is some combination of material properties and impact

pressure such that the solid responds as a viscous liquid, (ref. 2).

In essence, when the shear stress (t ) is sufficiently high the shear

max

£lflO 8SV; modulus goes to zero and rapid plastic flow will occur. Table I gives

the limiting particle velocity of the solid (

U

ps)

1 ^mi. above which this

phenomena will be present, for various liquid-solid combinations. This

type of deformation is commonly observed and has been used to qualitatively indicate the intensity and diameter of the impacting jet. From the

tabulated data one can qualitatively see the importance of both the solid

6
and the liquid physical properties.

When the solid typically fails in a brittle manner or when the liquid

impact velocity is insufficient to produce hydrodynamic deformation, then

some indication of the stress distribution in the solid is useful.

This can be done quite easily with the following assumptions: (1) only

elastic deformation occurs, (2) the impact pressure has a hemispherical

distribution with the maximum impact pressure determined by the RankineHugoniot relations, and (3) the impact pressure has a duration sufficiently long to allow a quasi-steady state to exist. The third assumption

is the most critical because the impact must last sufficiently long to

allow the stress wave to travel a distance such that the stress levels

are quite low, (see summary in ref. 2). Recent numerical and experimental efforts have shown that to a first approximation this assumption

is acceptable since the jet impact times have been estimated to be of the

order of 1 x 10 seconds. Figures 1 and 2 show the variation of the

maximum shear stress and one of the principal stresses, respectively,

and the manner in which they vary within the solid. The principal

results are as follows: (1) the impacted surface is under compression

and thus should not fail, (2) the maximum shear stress is within the

solid and below the impacted region and thus shear failure will occur beneath the surface, and (3) surface tensile stresses are present outside

the impacted region and thus brittle failure can start at the surface.

These same conclusions were reached when single impacts were experimentally studied for alumina, glass and plastic. Interesting enough, glass

has also been observed to have the above forms of failure when the impact

—8

duration was only of the order of 10 seconds. In any event, much more

study of the modes of material failure are necessary to understand the

mechanical aspects of the damage process.

In addition to mechanical damage, many investigations have shown that

electrochemistry can also have a significant role in the damage mechanism. As an example of the interrelation between cavitation and electrochemistry, it is instructive to review some particular experimental

7
results, (reference 3). In this experiment, a stationary specimen was located in a cavitating acoustic field. The subsurface properties: mosaic block size (i.e., subgrain size), microstress, and density

of dislocations, were determined by X-ray analysis for plain carbon

steel with a 8-16 y inch finish. These properties were compared for

"neutral", cathodic and anodic specimens. All of these measurements

were made in the so-called "incubation" period prior to any significant

material removal. The results are shown in Table 2. They demonstrated that the cathodic specimen was not measurably different from the

specimen having zero impressed current; both had significantly altered

subsurface properties due to cavitation. However, the anodic specimen

had very little subsurface property change due to cavitation but did

have a shorter incubation time. After the incubation period, the

anodic specimen maintained a relatively smooth surface whereas the other

two specimens developed a cratered surface. The application of the

impressed currents resulted in a lower surface microhardness compared

to the zero current condition, with the anodic specimen having the

smallest microhardness increase.

From many other tests it has been shown that materials with a low

corrosion resistance experience greater cavitation erosion when the

corrosivity of the liquid increases. Two of many possible explanations

would be that (1) the generation of microscopic galvanic action by material microstress gradients and (2) the continuous cathodic depolarization of existing oxygen concentration cells by high cavity collapse

induced velocities.

In summary, it can be seen that the physics associated with cavitation induced material damage is by no means completely understood.

Our understanding of the processes is just emerging and as more definitive information becomes available, our options in the design process

will become more apparent.

8
REFERENCES

1. "Hydrodynamic Cavitation & Some Considerations of the Influence of

Free-Gas Content", F.B. Peterson, Ninth Symposium on Naval Hydrodynamics, Paris, (August 1972).

2. "Some Considerations of Material Response Due to Liquid-Solid Impact",

F.B. Peterson, Journal of Fluids Engineering, vol. 1, pp 263-270,

(June 1973).

3. "Effect of Electric Current on Cavitation Erosion and on the Microscopic Substructure of Metal", A.F. Derendovski, M.K. Bologa and

K.K. Shal'nev, Dok, Akad, Nauk SSSR, vol. 83, pp. 1292-1294, (1968).

MAXIMUM SHEAR STRESS (Tm

/P,|

0.1 02 0.3

JHw OS.lT

FIGURE 1. Variation of maximum

Shear Stress (t ) with Radial and

m

Axial Distance from Point of

Impact.

SOLID LIQUID

T MAX

IN KILOBARS

,1Wum,t

IN METERS/SEC

Up, ^L,M1 T

Pb H 2 0 0.124 5.6 1.3

Pb NA 0.124 56 1.6

Pb Hg 0.124 5.6 8.0

AL

5456-H117 H 2 0 1 86 44 4 0 20

AL

5456-H117 NA 1 86 44.4 0 27

AL

5456-H117 Hg 1.86 44 4 1.2

TABLE 1. Hydrodynamic Deformation for Various Liquid-Solid

Combinations with Liquid Particle

Velocity Equal to 100 m/s.

9
FIGURE 2.

Variation of Princi

pal Stress (a./ )

if • and its Direction

(0) with Radial Dis

tance from Center

of Impact.

RADIAL DISTANCE FROM CENTER OF IMPACT

TABLE 2. Changes in Microscopic Substructure During Incubation Period

for ST 3 Steel (USSR).

(Specimen as (Specimen as Cathode) Anode)

Impressed current

Mosaic block size

Microstress

0

^ 30% dec.

*v» x 10 inc.

0.079 a/cm2

^ 30% dec.

*v> x 10 inc.

0.079 a/cm2

^ 4% inc.

no change

Density of dislocations ^ x 10^ inc. ^ x 10 1* inc. *v* x 2 inc.

Incubation time 105 min. 120 min. 75 min.

10
DISCUSSION

C. M. Preece, State University of New York, Stony Brook : You said

that the ductile materials seem to be quasi-viscous and that there

didn't seem to be any plastic deformation, that the deformation was

elastic. If there is a tremendous flow on the surface, it's permanent and therefore it must be plastic. In our own studies, we

have found an incredibly high dislocation density in this surface

which is more indicative that the deformation is plastic, not

elastic. Do you have any comments on this?

F. B. Peterson : Yes, I agree with you 100%. Apparently I gave

you the wrong impression. What I was trying to say is, that the

material deforms as if it's viscous. It is plastic deformation

without a doubt since, as we just pointed out, the deformation

is permanent. I think I said in the beginning that there were

smooth pits that remained after the impact, and that was the

particular situation I was describing. I was only talking about

the initial deformation of the solid. Certainly it is all based

upon the impact velocity, the particle velocity of the liquid,

and as the impact progresses, this velocity decreases. Then it

would slowly drift out of this so called hydrodynamic (my definition)

deformation and go into just elastic-plastic deformation.

J. Bonnin, Electricite de France : I must mention two points. You

noted that water is a unique liquid. I think that as far as the

nucleation is concerned, we cannot find any experimental liquid

in which we can find nuclei like in saline water. The second point

concerns particle size. You have mentioned particles as small as 0.2 ]M. I have found in the literature that Greenspan spoke of

particles as small as 0.01 or 0.02 ym.

F. B. Peterson : I believe you were agreeing with me that water is a unique fluid. You mentioned that you've seen references pertaining

to the measurement of particles which are as small as 0.01 to 0.02 ym.

For particles of that size, assuming that they are solid and unwetted

by water, a theoretical calculation indicates that it would take

something like 100 atmospheres of tension to cavitate.

J. Bonnin : But if such particles can be removed, negative pressures

up to some 200 atmospheres can be obtained.

F. B. Peterson : I agree. That was purely a theoretical calculation

if the cohesive bond is broken. But you're saying that the cohesive

bond doesn't have to be broken because there are solid particles in

the fluid. I was saying that if there are no solid particles, there

is an absolutely perfect fluid, in this case water.

11
G. 0. Ellis, Vickers Division, Sperry Rand Corporation : I am

interested in the van der Waals ' equation which indicates a large

tensile stress, and then a positive stress or a positive pressure.

When passing through the meniscus, there is a large negative pressure that indicates that there can't be cavitation, but then there

is a significant positive pressure, I believe about 750 atmospheres.

I would assume that at a certain bubble size there would be a high

positive pressure. There are theories that say that water has a crystalline structure for a few molecules, and so it may not be a bubble that forms, but rather a rip or a tear. I wonder what you

think of that theory?

F. B. Peterson : With regard to the van der Waals' forces, the

consideration that I was really talking about was the energy required

to separate the fluid, which would mean moving the molecules about

3-10 angstroms apart in order to overcome this cohesive force. You

are talking about a gas-liquid interface in the case of a bubble,

and I haven't thought about it from that standpoint. I have thought

about it more from the standpoint of surface tension, maintaining a pure surface, etc. With a surfactant, there would certainly be

changes in these forces that you are talking about. I suspect it

would be fairly difficult, at least in water, to sustain a pure

vapor-liquid interface without having some surfactant there. The

condition I was talking about of a solid coming out of the melt is

an example of the breaking of the interfacial bond that existed

between the liquid molecule and the solid, the low bond strength

that existed there. What was happening was not microscopic, it

was something that was visible. But if you want to talk about it

as a tear in the fluid, than I guess that would be okay.

12
ROLE OF PHYSICAL PROPERTIES OF LIQUIDS IN CAVITATION EROSION

A. Thiruvengadam

Professor of Mechanical Engineering

The Catholic University of America

Washington, D. C. 20017

Cavitation erosion is a serious engineering problem in hydraulic

turbines, pumps, valves, control devices, hydrofoils, propellers, bearings, diesel liners, aircraft engines, torque convertors, acoustic devices and processing and cleaning equipment. Basic research on bubble

growth and collapse, material erosion and environmental influences has

aided the design engineer in the proper design and material selection

process. Various screening tests provide the relative erosion strength

of candidate materials and protection techniques. There were cases in

which the erosion intensities were too severe to be protected by existing materials and protecting techniques. In such cases operational

limitations and redesign become necessary. Model testing and redesign

would greatly help the practical engineers. The properties of liquids

play an important role in the development of scaling laws and model

testing techniques. They influence fluid dynamics, bubble dynamics and

environmental effects.

Experimental and theoretical research has been done on many engineering liquids including water, seawater, various lubricants, coolant

additives such as glycols and alcohols and liquid metals including

13
mercury, sodium, potassium, lithium, and cesium. Physical properties

studied include density, vapor pressure, surface tension, gas content,

sound speed and viscosity. Some of these properties have been experimentally varied by controlling temperature and pressure of the liquid.

A summary of all the experimental results obtained by various investigators is available in reference (1). The erosion rate generally increases with increasing temperature of the test liquid, reaches a maximum and decreases as the liquid temperature approaches the boiling

point. While this is true in most of the liquids tested, there are a

few exceptions. Commercially pure titanium (100-A) rested in liquid

sodium in the temperature range from 400°F to 1000°F shows the above

trend; whereas 316 stainless steel tested in the same temperature range

shows a monotonically decreasing erosion rate with increasing temperature of the sodium. Any attempt to explain the influence of ambient

temperature on erosion rate must take into account the effect of temperature on the energy absorbing characteristics of the material.

There are some data available on the erosion rate of 316 stainless steel

in mercury up to 150°c. The erosion rates in liquid sodium and in mer-n

cury are comparable around 150° c.

Nowotny (2) compared the erosion rates in water, aniline, phenol,

toluene and amyl alcohol and concluded that there existed a correlation

between the surface tension of the liquids and the erosion rates. More

recent experiments by Plesset (3) and Thiruvengadam (1) confirm this

Ik
conclusion. However, Wilson and Graham (4) studied the effects of density, compressibility, vapor pressure, and viscosity in lubricating oils

and concluded that the acoustic impedence (the product of density and

speed of sound) played a major role. Thiruvengadam (1) showed that the

dissolved air content was also important. More recently, Plesset (5)

conducted cavitation erosion tests on aluminum in several mixtures of

non-aqueous liquids and water, and found that erosion rates decreased

monotonically as a function of concentrations, except in one case where

a minimum was observed. Plesset concluded that pure aluminum was chemically less active in non-aqueous liquids. Wilson and Graham (4) as

well as Thiruvengadam (1) showed that the viscosity of the liquid did

not have much effect on erosion rate.

The above results indicated that a systematic examination of the

role of physical properties on cavitation erosion was needed. The math^-

ematical analysis for examining the effects of physical properties of

liquid on cavitation erosion can be traced to the work of Rayleigh (6).

He obtained infinite bubble wall velocities and collapse pressures for

an empty bubble collapsing in an inviscid incompressible liquid; he

then made the assumption that a permanent gas obeying Boyle's law was

inside the bubble, which was compressed isothermal ly during collapse.

Furthermore, in his conclusions, he suggested that a more realistic theory should take into account the compressibility of the liquid. Herring

(7), Trilling (8), and Hickling and Plesset (9) considered the effect

15
of compressibility while Gilmore's (10) investigation included a discussion of the effects of viscosity and surface tension. Ivany and

Hammitt (11) combined the effects of viscosity and surface tension with

compressibility during the collapse of empty and gas-filled bubbles.

The effect of surface tension was brought to attention by Plesset (12)

and Noltingk and Neppiras (13), whereas the viscosity of the liquid was

investigated in detail by Poritsky (14) and Ivany and Hammitt (11) .

In general no more than three physical properties of the liquid was

considered in their analyses.

It was evident that a systematic investigation of the role of physical properties in cavitation erosion was needed. With this objective,

Tung and Thiruvengadam (15) conducted both theoretical and experimental

studies. The analytical study was based on the idea that the intensity

of erosion is related to the intensity of bubble collapse forces (16).

Considering both spherical and non-spherical collapse, this analysis

leads to a relation between the rate of erosion and the product of the

collapse pressure, frequency of collapse and the radius of the bubble.

The maximum collapse pressure is related to the surrounding pressure,

the gas content inside the bubble, the sound speed and density of the

liquid. The maximum size of the bubble is related to the growth time

available, the pressure difference between the inside and outside of the

bubble and density of the liquid. The frequency of growth and collapse

of the bubbles are related to nuclei size distribution, surface tension,

velocity, pressure and size of the system. When all these parameters

16
have been examined, and expressed in terms of physical properties of the

test liquid, a relationship governing the intensity of bubble collapse

may be derived. ' This relationship closely agrees with many different

experimental results. For example, the decrease in erosion rates in wa-^

ter-alcohol mixtures is related to the change in the physical properties

of the water-alcohol mixtures as predicted by the analytical relationship.

The dependence of erosion rates on the ambient temperature of water

can also be predicted quantitatively with this relationship. The assumption that the gas inside the bubble is compressed adiabatically during

collapse gives a better agreement with experiments than the assumption

that the gas is isothermally compressed as assumed by Rayleigh. Calculations for the eleven liquids tested by Wilson and Graham agree closely

with experimental data. Acoustic impedence is an important liquid parameter that governs the erosion intensity in vibratory devices. This

investigation reveals that the major physical properties of liquids governing the intensity of erosion include density, sound speed, surface

tension, vapor pressure, gas content and nuclei distribution.

Although these results are generally useful in studying the erosion

potential of various engineering liquids and in selecting scaling parameters for model studies, it is only a first step in the understanding

of the complex processes involved in cavitation erosion. For example,

the measurement and characterization of nuclei distribution is needed

17
in these studies. , Another important area is the measurement of the

sound speed and density in bubble-water mixtures. Current estimates

show that the sound speed can be as low as 100 fps in air -water mixtures. It is evident that much more remains to be accomplished for an

overall understanding of the phenomenon of cavitation erosion.

ACKNOWLEDGEMENT

This research was supported by the U. S. Office of Naval Research

under Contract No. NOOOlU -67 -A -0377-0008.

18
References

1. A. Thiruvengadam, "On Modeling Cavitation Damage", J. Ship Research,

Vol. 13, pp. 220-233, 1969.

2. H. Nowotny, "Destruction of Materials by Cavitation", VDI Verlag,

Berlin, 1942.

3. M. S. Plesset, "Cavitation Erosion in Non -Aqueous Liquids", J. Basic

Engr., Trans. ASME, Vol. 92, Series D, pp. 807-818, 1970.

4. R. W. Wilson and R. Graham, "Cavitation of Metal Surfaces in Contact

with Lubricants", Conference on Lubrication and Wear, Institution of

Mechanical Engineers, Paper 83, Oct. 1957.

5. M. S. Plesset, "Temperature Effects in Cavitation Erosion", ASME

Winter Meeting, Paper No. 71-WA/FE-30, 1971.

6. Lord Rayleigh, "On the Pressure Developed in a Liquid During the

Collapse of a Spherical Cavity", Phil. Mag. 34, Series 4, pp. 94-

98, 1917.

7. C. Herring, "Theory of the Pulsations of the Gas Bubble Produced by

an Underwater Explosion", Columbia University, NDRC Rep. C-4-SR 10-

D10, 1941.

8. L. Trilling, "The Collapse and Rebound of a Gas Bubble", J. Applied

Phys., Vol. 23, pp. 14-17, 1952.

9. R. Hickling and M. S. Plesset, "Collapse and Rebound of a Spherical

Bubble in Water", Phys. of Fluids, Vol. 7, pp. 7-14, 1964.

10. F. R. Gilmore, "The Growth and Collapse of a Spherical Bubble in a Viscous Compressible Liquid", Calif. Inst, of Tech. Hydrodynamics,

Lab. Rep. 26-4, 1952.

11. R. D. Ivany and F. G. Hammitt, "Cavitation Bubble Collapse in Viscous, Compressible Liquids Numerical Analysis", Trans. ASME, Vol. 87,

Series D, J. Basic Engr., pp. 977-985, 1965.

12. M. S. Plesset, "Bubble Dynamics", In Cavitation in Real Fluids, pp.

1-18, Elserier Publishing Co., Amsterdam, 1964.

13. B. E. Noltingk and E. A. Neppiras, "Cavitation Produced by Ultrasonics", Proc. Phys. Soc. 63. B, pp. 674-685, 1950.

19
14. H. Poritsky, "The Collapse or Growth of a Spherical Bubble or Cavity in Viscous Fluid", Proc. First U. S. National Congress on Applied Mechanics (&SME) , pp. 813-831, 1952.

15. S. Tung and A. Thiruvengadam, "The Role of Physical Properties of

Liquids in Cavitation Erosion", Seventh Southeastern Conference on Theoretical and Applied Mechanics, 1973.

16. A. Thiruvengadam, "Scaling Laws for Cavitation Erosion", Proc. Symp

on Flow of Water at High Speeds, The International Union of Theoretical and Applied Mechanics, Leningrad, U.S.S.R., June 1971.

-

-

20
DISCUSSION

F. B. Peterson, Naval Ship Research § Development Center: I wonder

if you meant to imply by one of your earlier slides where you used

distilled water and various solutions of sodium chloride that cor- rosion was not a factor? Were you saying that corrosion was not

an important aspect in the material loss? For a wide variety of

flows, at least those that exist in systems of the Navy, erosion

is really a major problem when there are cavities which are basically

fixed to the surface as if there would be a separated flow regime.

An example would be cavitation that exists between the intersections

of the strut and a foil in a hydrofoil system, or on blades or propellers, or in pumps. I wonder if the Rayleigh calculations are

the most appropriate ones when the spherical bubbles you're talking

about are generated from the break up of a large vapor cavity?

Lastly, would the Rayleigh equations really be valid for collapse

when there is the strong jetting mechanism?

A. Thiruvengadam: As far as the influence of corrosion is concerned,

it is not just the liquid alone that produces the corrosion results

of material -liquid combinations. All we are saying is that if a commercially pure aluminum (99.99%) is exposed to salt water for

15-30 minutes, the corrosion influence is small compared to the erosion

influence. This is the point we were talking about in terms of

non-aqueous liquids. Whereas, if steel is cavitated for about 3 hours,

there can be a factor of 10 difference depending upon the amplitude.

Using lower amplitudes results in lower intensities, longer times,

and the corrosion influence dominates the erosion influence. Coming

to the second aspect of your question, we know in Navy systems why

bubbles collapse when the cavity is slightly developed. At a^, the

so-called degree of cavitation is implied by saying there is a fixed

cavity at which the erosion is a maximum. Beyond the inception stage

these bubbles will fully develop, i.e., reach maximum size, and then

collapse. At that time, the erosion intensity is a maximum. If it

grows further down, the pressure must be reduced, or the frequency

has to be changed. These are two of the parameters, all of which

change, that go into making up the intensity of erosion. They are

all dependent upon the a value. Now, can we assume Rayleigh bubble

growth? Here again, it's an extremely complex flow field, as you

pointed out. We assume that the bubbles are formed near the inception

point and then are carried away by the flow. Dr. Hammitt has con- ducted some experiments where he photographed some bubbles that were

supposed to be carried at free stream velocity values. Plesset

conducted some analytical work where he traced the bubble growth along

the guiding boundary and used the Rayleigh distribution to predict

21
the maximum size, growth time, collapse time, etc. I agree that

it is a very complex area. The Rayleigh equations cannot really

be used in a 100% practical sense, but that's the place to start

in this kind of analysis.

J. P. Tullis, Colorado State University: Since the erosion rate

in a vibratory system is orders of magnitude faster than in a flowing system, are the relationships that are derived from this

vast amount of data directly applicable, or can we make the transition to flowing systems which are, of course, the practical problems?

A. Thiruvengadam: This has been one of our objectives in doing the

so-called modeling techniques, counting techniques, etc. The erosion

intensity in the field need not necessarily be lower than in the

laboratory apparatus. We can give examples in which the erosion

intensities can be 100 times more severe than in the laboratory

apparatus. We have all been concerned with how to extrapolate

laboratory data to the field in a practical engineering system

to guide designers.

C. M. Preece, State University of New York, Stony Brook : I want

to comment on Dr. Peterson's first question about the corrosion.

It's about time we all stopped thinking of aluminum as not being

corroded by distilled water. In a dynamic system where passivation

is prevented, distilled water is the corrosive environment for

aluminum, and therefore little difference would be expected between

the behavior of a sodium chloride solution and pure distilled water.

F. G. Hammitt, University of Michigan : In your calculation of the

effects of temperature on damage rate, how can you apply Rayleigh 's model which as far as I know has no way to take into account the

different temperature effects? There is no thermodynamic effect

in Rayleigh' s model. How did you take that into account?

A. Thiruvengadam: We got the calculations through the bubble growth

times. The temperature comes in ok. When the temperature varies,

the physical properties vary, density varies, sound speed varies,

surface tension varies, vapor pressure varies, dissolved gas pressure

varies. For each temperature, we find out what these values are. One of the reasons we used water and water-alcohol systems is that some of these values are readily available in the International

Critical Tables.

22
THE ROLE OF MECHANICAL PROPERTIES IN

CAVITATION EROSION RESISTANCE

G. C. Gould

Materials and Processes Laboratory

General Electric Company

Schenectady, New York 12301

There has been a long standing need for a correlation between the

cavitation erosion resistance of materials and some of their common

mechanical properties so that a prediction of erosion resistance would

be possible from readily available mechanical property data. In addition to being of great practical value a correlation of this type

would infer a mechanism of cavitation erosion damage or inherent material

erosion resistance that currently is lacking.

In order that correlations of erosion resistance and mechanical

properties of materials be attempted the erosion resistance of materials

must first be measured. There are several ways of doing this experimentally:

1) Vibratory Apparatus

a) Piezioelectric

b) Magnetostrictive

2) Venturi Tunnels

3) Rotating Discs (with holes)

4) Impingement Apparatus

a) Whirling Arm

b) Stationary Target

23
By far the most common method of testing cavitation erosion resistance of materials is the vibratory cavitation probe. In this device the

specimen is placed on the probe tip and submerged in water (or any other

liquid), is vibrated at a high frequency, 20,000 Hertz, and cavitation

is induced in the liquid adjacent the specimen tip. The cavitation

causes damage on the specimen and the damage is normally monitored by

measuring weight loss of the specimen. The second most popular method

of measuring erosion resistance is the whirling arm device. This device

moves the specimen through a relatively stationary field of droplets at

a high velocity and the resulting impact of droplets on the specimentarget cause erosion damage and this damage is also measured as weight

loss

.

Hobbs^ has shown that the erosion resistance of materials is relatively the same regardless of which of these two techniques is used to

rank materials. For the sake of this discussion it is assumed that cavitation erosion damage and liquid impingement damage are the same general

phenomanon and data gathered by either technique is useable for correlation attempts. There is significant scatter among the various cavitation tests when they are compared to one another as in the ASTM Round

Robin tests reported by F. G. Hammitt.^

The first and most common correlation of erosion resistance and

mechanical properties attempted has been between hardness and erosion

resistance; however, this has not been successful when a wide range of

(3) materials has been analyzed. 1 A better correlation has been established using the strain energy of the material and correlations based on

2k
various forms of this quantity have been put forth by Thi ruvengadam(4)

,

Hobbs^, Backstrom^, and Thomas and Brunton^. The correlation found

with erosion resistance and strain energy terms is better than that found

with hardness; however, there are notable exceptions.

Metallographic investigations v ' of the earliest stages of cavitation erosion damage of metallic materials show plastic deformation occurring during the incubation period and increasing until cracks form

and metal fragments are lost. This observation supports, in general, a

hypothesis that takes into account the total energy for fracture as an

important material variable for predicting erosion resistance.

The attempts to correlate erosion resistance with mechanical properties involved the use of parameters such as work done to cause fracture (Proof Stress - 2/3) (Tensile strength - Proof Stress) X Elongation,

Proof Resilience (1/2 (Proof Stress)2/Elastic Modulus) and Ultimate

Resilience (1/2 Tensile Strength) /Elastic Modulus), all of which try to

take in£o account in some way the energy absorbing capacity of the ma-

(5) terial. Hobbs v ' has shown that the erosion resistance versus ultimate

resilience correlation is the best of this lot, but again the correlation

becomes poorer the greater the amount of data used.

Considering the absorption of energy the important role of a material in resisting erosion by cavitation the idea of erosion resistance

depending on strain energy during deformation of a material has been advanced by Brunton,^ among others. Again, the correlation works well

with a limited set of data but as more kinds of material are added the

correlation becomes poorer. In particular cobalt appears, in Brunton's

25
work, to be far more erosion resistant than its strain energy would indicate. This observation is of more than passing interest since some of

the most erosion resistant materials are the Stellite alloys which are

cobalt based alloys.

Investigation of the Stellite alloys has been reported^'^ and in

these investigations a deformation induced phase transformation has been

found and the transformation of FCC to HCP has been cited as the energy

absorbing mechanism that accounts for the superior erosion resistance

of the Stellite alloys.

Work in the field of explaining and predicting the erosion resistance of materials is continuing, and needs to continue. There is no way,

now, of predicting the erosion resistance of materials from their mechanical properties.

REFERENCES

1. Hobbs, J. M. /'Practical Aspects of Cavitation", Phil. Trans. Royal Soc .

of London , Vol. 260, 1966.

2. Hammitt, F.G.,"ASTM Round Robin Test with Vibratory Cavitation and

Liquid Impact Facilities", Materials and Research Standards, Vol. 10,

1970.

3. Gould, G.C.,"Some Observations on Erosion by Cavitation and Impingement", ASTM STP-474 , 1970.

4. Thiruvengadam, A., "A Unified Theory of Cavitation Damage", AS ME

Journal of Basic Engineering , pp. 365-376, September 1963.

5. Hobbs, J.M., "Experience With a 20-KC Cavitation Test", Erosion by

I

26
Cavitation and Impingement , ASTM STP-408 , 1967.

6. Backstrom, T.E., "A Suggested Metallurgical Parameter in Alloy Selection for Cavitation Resistance", United States Department of the

Interior , Bureau of Reclamation Report No. ChE-72 , December 1967.

7. Thomas, G.P. and Brunton, J.G., "Drop Impingement Erosion of Metals",

Proc. Royal Soc. London A314 , pp. 549-565, 1970.

8. Gould, G.C., "Cavitation Erosion of Stellite and Other Metallic

Materials", Proceedings of the Third International Conference on Rain

Erosion and Associated Phenomena , RAE August 1970.

9. Woodford, D.A. , "Cavitation Induced Phase Transformations on Alloys",

Met. Trans . Vol. 3, May 1972.

27
DISCUSSION

R. E. Maringer, Battelle, Columbus Laboratories : Has any attempt

been made to associate the initial period of cavitation with

residual stresses in the surface due to machining? These penetrate

depths, depending upon the hardness, of 5, 10, 20 mils, which sort

of correspond with the change of rate of cavitation. These are ignored in most mechanical property tests, but when dealing with

something that is so surface sensitive, they have to be watched

particularly.

G. C. Gould : Having just been burned very badly with that in an entirely different program, I can't agree with you more. You're

absolutely right. I think, to some extent, they have been ignored.

I know in my own instance I've done some electro-chemical polishing

of specimens to try to get around that, and I am not certain of how

well I've done it. It would appear in materials like austenitic

stainless steel, as an example, and probably in stellite or in cobalt,

that the incubation period might be lengthened or shortened by how

much residual stress, or inherent plastic deformation, is there

'because of sample preparation. During the incubation period, the

material is severely deformed, and then the process of chipping it

away proceeds. So I would expect that it has residual stresses, or

that specimen preparation has its biggest influence on the length

of the incubation period, and not so much on the subsequent, what

we like to call rather glibly, steady state.

F. G. Hammitt, University of Michigan : I have some confirmation of

what you just said. We did some experiments where specimens were

held under tension and also under compression near the yield point

or beyond it. For compression, it seemed to inhibit the initial

pitting and thus did, in fact, increase the nucleation period.

Tension didn't seem to have much effect, at least not a reproducible

effect. Neither of them seem to have very much effect in the developed

damage portion, much as you summarized there. Another point I want

to mention is that about 10 years ago, there was some work reported

by Plesset where x-ray diffraction techniques were used to examine

the surface at the very beginning of cavitation. It was found that

there were some distortions in the surface long before any noticable

weight loss could be found. I don't think that work has been

repeated. But it's a good approach, I think.

28
C. M. Preece, State University of New York, Stony Brook : I am not

certain that the surface effect is necessarily the important one,

because our experiments indicate that it's a bulk phenomenon far more

than a surface phenomenon. As Professor Hammitt said, experiments

indicated that a bulk stress, not a surface stress, was being applied.

We have, in fact, repeated some of the work that Plesset did on single

crystals of aluminum. One of the things that surprised us was that

a single crystal of aluminum which has nice, sharp diffraction patterns

prior to exposure to erosion, has streaked diffraction spots from the

front face of the sample after exposure, but so do the ones from the

back surface. The deformation is definitely a bulk phenomenon.

Furthermore, in a cobalt sample with a face-centered cubic structure,

it was found that both the front and the back of a several millimeter

thick sample is transformed after about a 1/2 hour exposure to

cavitation without any detectable weight loss. So I think it's more

the state of the bulk of the material rather than the surface which

may be important.

G. C. Gould : I can't let that go uncontested. In your work you're

using a specimen on an anvil, or some such thing, and a probe with

a non-wearable tip. You're irradiating it, if you will, with a cavitation environment, and that's what you observe. I don't dispute

that, but if I take the stellite specimens which are the tips and

are not in any way able to be irradiated in the back and held reasonably

rigidly in this titanium mechanical transformer, I find by means of

x-ray diffraction a transformed layer on the surface. This electropolishes off with just a few milligrams of dissolution, a non-deformation

kind of dissolution. It would appear under these circumstances that

the effect of the cavitation is rather minimal as far as penetrating

into the depth of the specimen is concerned.

C. M. Preece : Maybe you are transforming the sample into a non- equilibrium phase. I think if you are reflecting the shock pulse, or whatever it is, from the reverse side of the sample, that you may be

undoing what you have already done. It has been shown in shock

deformation that the transformation is reversible, in steels at any

rate, and I think perhaps this is one explanation. In our situation

where we're going in pure cobalt from an unstable face-centered cubic

to a stable HCP phase, it may remain undone. In your case, it may

just be the stability of the two phases which is determining it.

29
E. E. Klaus, The Pennsylvania State University : There was some suggestion earlier that there might be a chemical effect in cavitation erosion. I wonder whether you have identified any chemically

reactive films on the solid surface?

G. C. Gould : I have not. I have done some work patterned directly

after Plesset's in which, with time, instead of just turning on a given level of cavitation intensity, it has been pulsed. Now the

specimen enjoys ten fold greater exposure to the environment during

the same total amount of damaging cavitation, in the system that

I have looked at doing this, namely copper in water, which is a readily oxidizable material, there seems to be no difference. There

seems to be no difference whether the water is oxygenated or deoxygenated by purging nitrogen through it. That doesn't rule out

the possibility that corrosion is going on, but the amount of corrosion

going on is quite minimal with respect to the overall damage rate,

which in copper is very high. I guess my experiment isn't a real

critical one. Plesset has shown with plain carbon steel in sodium

chloride solutions that there is clearly a contributing corrosion

effect. I've done a lot of work with 12 chromium steels, and I have not

found a lot of corrosion products. Maybe I shouldn't have expected

to, and maybe that material passivates so fast that the corrosion in

that case is unimportant.

J. E. Stern, NASA, Goddard Space Flight Center : In regard to the

electro-polishing of your specimens before tests, what are the

possibilities that you might have encountered some hydrogen embrittle

ment that might have had an effect on the behavior of the surface?

G. C. Gould : The thought of hydrogen embrittlement went through my

mind when Frank Peterson mentioned the Russian work. All I can say is,

it's a possibility. I don't know the effect on a cobalt base material

when it's permeated with hydrogen. Being face-centered cubic, it

would probably be much less likely to be embrittled than would a bodycentered cubic material, but I really don't know the answer.

30
RECENT THEORIES OF CAVITATION DAMAGE INCLUDING

NON-SYMMETRICAL BUBBLE COLLAPSE EFFECTS

Frederick G. Hammitt

Department of Mechanical Engineering

The University of Michigan

Ann Arbor, Michigan

Recent theories of cavitation damage mechanisms will be discussed

in the present paper as opposed to the classical theories on that subject,

originated primarily by Rayleigh in 1917 (1), who considered the collapse

of an empty spherical bubble initially at rest in a static, incompressible,

inviscid, infinite liquid. It was assumed in the various studies based

on this initial work that at least the mechanical portion of cavitation

damage (as opposed to chemical effects) resulted from shock waves

propagated through the liquid by the collapse of cavitation bubbles under

the ideal fluid and spherical symmetry assumed by Rayleigh.

A more realistic approach to the problem requires primarily a consideration of the non-symmetries inherent in most actual cavitation

collapse situations. Photographic evidence, dating primarily from the

work of Naude and Ellis (2), shows that the actual collapse of bubbles

near a symmetry-destroying feature such as a nearby wall results in a toroidal-like collapse, with final generation of a liquid microjet oriented toward the wall. However, this has since been substantiated by

numerous photographic studies, and also theoretical studies of varying

and increasing complexity (3,4, e.g.). Present photographic and

numerically-computed evidence indicates that the microjet velocity is

31
only marginally sufficient to be the sole cause of the observed cavitation pitting in strong materials.

Relatively recent numerical calculations including numerous real

fluid effects (5,6, e.g.) indicate that the shock wave intensity emitted

during collapse is not likely to be strong enough to be damaging to most

materials, even if spherical symmetry were maintained through a very

large volume ratio, and if the bubble were touching the wall at the start

of collapse. Of course, spherical symmetry is impossible for the

collapse of a bubble in such a position, but the shock wave intensity is

in general of maximum intensity for a given volume ratio of collapse,

if spherical symmetry is maintained. However, the numerical calculations go on to show that if a "rebound" occurs such that the bubble grows

again after collapsing, the event is similar to a micro-explosion (rather

than implosion as for collapse), and in this case the radiated shock-waves

are much stronger, so that damage through their impingement on a neighboring wall is much more likely. Photographs by various investigators, including Kling of our own laboratory (7,8, e.g.) show that nonsymmetrical collapses with rebound occur almost exclusively in a real

flowing system. Hence it appears to this writer that actual damage is

usually a result of a combination of the impact effect of the microjet and

the shock-wave pressures generated by bubble rebounds. There is much

more confirmatory evidence than that which can be covered in this

relatively brief paper. However, much of the general applicable material

is covered in the book Cavitation (9), as well as an extensive pertinent

bibliography.

A final pertinent point is the photographically demonstrated and

theoretically justified fact that the bubble centroid during collapse is

32
strongly attracted toward a solid wall, and that the microjet is then

directed toward the wall, provided other influences such as velocity or

pressure gradients are not sufficient to overcome the wall-attraction

effect. For these reasons, a relatively strong sorting mechanism

exists preventing the great majority of bubble collapses in the vicinity

of a wall from reaching damaging intensity. This is confirmed experimentally. On the other hand, a collapse near a free (or very flexible)

surface, affects the bubble behavior in a precisely opposite manner.

The bubble is repelled from the flexible surface and the jet is directed

away. This also is observed photographically (10, e.g.) and is also

expected on theoretical grounds. It may explain the surprisingly good

cavitation resistance of various elastomeric materials.

References

1. Lord Rayleigh, "On the Pressure Developed in a Liquid during the

Collapse of a Spherical Cavity", Phil. Mag. , 34, 200, 1917, 94-98.

2. C. F. Naude and A. T. Ellis, "On the Mechanism of Cavitation

Damage by Non-Hemispherical Cavities Collapsing in Contact with

a Solid Boundary", Trans. ASME, J. Basic Engr . , 83 , D, 4, 1961,

648-65 6.

3. M. S. Pies set and R. B. Chapman, "Collapse of an Initially Spherical Vapor Cavity in the Neighborhood of a Solid Boundary", J_. Fluid Mech. 47, 2, May 1971, 283.

4. T. M. Mitchell and F. G. Hammitt, "Asymmetric Cavitation Bubble

Collapse", Trans. ASME, J. Fluids Engr. 95, 1, 1973, 29-37.

5. R. Hickling and M. Plesset, "Collapse and Rebound of a Spherical

Bubble in Water", Phys. of Fluids , 7, 1, 1964, 7-14.

6. R. D. Ivany and F. G. Hammitt, "Cavitation Bubble Collapse in

Viscous, Compressible Liquids, Numerical Analysis ", Trans

.

ASME, J. Basic Engr. , 87, D. , 4, 1965, 977-985.

33
7. C. L. Kling and F. G. Hammitt, "A Photographic Study of SparkInduced Cavitation Bubble Collapse", Trans . AS ME, J. Basic

Engr . , 94 , D, 4, 1972, 825-833. ~

8. C. L. Kling, F. G. Hammitt, T. M. Mitchell, E. E. Timm,

"Bubble Collapse Near a Wall in Flowing System", 16 mm motion

picture (silent) ASME Film Library, 1969.

9. R. B. Knapp, J. W. Daily, F. G. Hammitt, Cavitation, McGrawHill, 1970.

10. E. E. Timm and F. G. Hammitt, "Bubble Collapse Adjacent to a Rigid Wall, Flexible Wall, and a Second Bubble", 1971 ASME

Cavitation Forum, 18-20.

I
DISCUSSION

J. Bonnin, Electricite de France : You noted that when the cavitation

parameter is reduced, there. is sonic inception followed closely by

visible inception. These are very close together. This is true,

I quite agree with you, but only in cold liquids. You showed a figure with curves with visible inception, first the marriage, sonic

inception, etc. We can observe with liquid carbon dioxide that we

could have cavitation inception, and even develop some cavitation

without any noise. But it was, I must say, in carbon dioxide and

very close to its critical point.

F. G. Hammitt : We have generally found that if you look carefully

enough, when you can hear something you can see something, or vice

versa, at least in our venturi with water. But I have no experience

with carbon dioxide. We've done some work with liquid metals, but

then nothing can be seen. Sonics must be relied on.

R. M. Whittier, Endevco : Have you done experiments using ultrasonic

stress wave sensors and things like this?

F. G. Hammitt: No.

35

SESSION II

Chairman: R. S. Miller

Office of Naval Research

SOME PRACTICAL EXAMPLES OF CAVITATION EROSION

AND THEIR PREVENTION

Andrew F. Conn

HYDRO NAUTICS , Incorporated

Laurel, Maryland 20810

PREFACE

The concepts and approaches to the prevention of cavitation damage which are outlined in this abstract are almost

entirely those presented in the two recent review articles

by A. Thiruvengadam (1, 2). As these reviews contain such a

wealth of detail and references about this complex subject,

and as Dr. Thiruvengadam taught this author what little he

knows about cavitation, no attempt will be made here to repeat this information in any detail. The reader is strongly

urged to obtain Reference 1 if a complete review of this

subject is desired.

INTRODUCTION

This presentation surveys the problem of failures caused

by cavitation erosion from an engineering point-of -view,

i.e.., with de-emphasis of on-going "state-of-the-art" research controversies, in an attempt to provide the designer

and engineer with some tools which can be used to: a) assess

the magnitude of his cavitation erosion problem, and b) rationally select from one or more alternative solutions.

39
The concepts of intensity of erosion, erosion strength

and the time dependence of erosion rate are described; and

the relation of these parameters to system variables such as

pressure and velocity,, and to the properties of materials

are discussed. Using several examples of actual cavitation

erosion, methods of prevention and their limitations are examined .

CAVITATION EROSION PARAMETERS

Any assessment of damage from cavitation erosion requires,

at least, some quantitative knowledge of the amount of material removed and the exposure time in order to estimate the

rate at which erosion has proceeded. Although the time history of erosion typically has four periods, namely: incubation, acceleration, deceleration, and steady, one may use,

particularly in more severe cases, an averaged rate based on

the overall elapsed time.

The severity of cavitation may be characterized by the

intensity of erosion, I, which has the dimensions of power

per unit area. Physically, this intensity parameter represents the rate at which energy is being absorbed by the

eroding material, and it is highly dependent on the velocity,

pressure, and temperature in the system, and on various fluid

properties such as viscosity and surface tension. Minor design variations can greatly influence the value of I in a

system, causing severe localized erosion.

ho
Although in general the harder and stronger materials are

more resistant to cavitation erosion, the search continues

for a single parameter, easily measured or inferred from

other material properties, to characterize this resistance.

One finds, for limited classes of materials, that good correlations can be made between erosion resistance and almost any

parameter such as hardness, strain energy, ultimate strength,

or "ultimate resilience". However, the high frequency, repetitive nature of the cavitation erosion process, and the

inherent variations in the response of materials to this

type of energy input, makes one wonder if a universal correlation factor is worth seeking.

Several investigators have suggested the use of some type

of relative erosion resistance scale, based on the results

of a standardized test procedure. Among these are Heymann's

"normalized erosion' resistance ", N

g (3) and Thiruvengadam '

s

"erosion strength", S

e (4). Thus, without requiring a full

understanding of the role, if any, of other material properties, one may subject a variety of materials to one of the

screening tests, and acquire what is needed for design purposes .

It may be shown (4) that the rate of erosion, r , is related to the intensity of erosion, I , and the erosion

strength, S

g

, by: r = i/S j and this relation may be used

to compute any one of these parameters after the other two

have been determined.

hi
ASSESSMENT AND FAILURE PREVENTION

Faced with a system that has experienced damage from cavitation erosion, and concerned with preventing a failure either by fracture due to weakening of the material or hydrodynamic degradation below a functional level, the designer

must first assess the seriousness of the problem, and then

select the most effective approach to eliminate or at least

minimize the damage.

In the previous section of this abstract, the basis for

assessing the magnitude of the problem, namely, the intensity of erosion, I, is outlined. The rate of erosion must be

estimated, and with the erosion strength, S

g

, of the material being eroded, one can then make an estimate of intensity.

If the S

g

is unknown, one of the screening techniques such

as the vibratory test, the rotating disk test, the venturi

test, or a jet impact test can be used on samples of candidate materials to provide this evaluation.

If the intensity of erosion is found to be around 10

_1

watt/meter2or less, then one of the following methods of

protection should be effective: substitution of a more resistant material, either for the entire component or at

critical areas in the form of either inlays or overlays; use

of elastomeric coatings; cathodic protection; or air injection.

In the intermediate intensity range from about 10 1 to 1

w/m2

, it may be possible, in addition to one or more of the

k2
above, to sufficiently reduce the rate of damage by imposing

operational limitations on velocity, power, or pumping capacity. In many cases it has been found that the intensity of

erosion is proportional to the velocity raised to an exponent of six or greater. Thus, a velocity reduction of 10$

could halve I; and 20$ less speed could reduce I by almost

75$, which would quadruple the operating lifetime.

Intensities of erosion above 1 w/m2 will usually require

major design changes. Provision for air vents, or severe

limitations in operation may prolong the system lifetime,

but hydrodynamic redesign or frequent periodic maintenance

are probably the only practical solution in these cases.

It should be emphasized that, for non-aqueous fluids, and

with other than ambient conditions the screening tests must

be conducted so as to approximate the actual service conditions as closely as possible. Much research remains to be

done before predictions of material erosion can be made from

simple laboratory tests. Thus, for severe situations, the

expected pressures, temperatures, times, and corrosive factors should be imposed during the screening tests.

CONCLUDING REMARKS

Much remains to be learned about the basic nature of the

cavitation erosion phenomenon. Research is actively underway throughout the world (see Ref.5) on aspects such as fluid

properties, bubble dynamics, size scaling, and the role of

various material parameters in the resistance to erosion.

Studies into the failure of coating-adhesive-substrate
systems continue to raise more unanswered questions, and

trial and error is still the main tool for the hydrodynamicist faced with a badly eroding propeller or pump. However,

despite what we still need to learn, the approach outlined

in this presentation allows the designer to understand the

magnitude of his cavitation erosion damage in terms which

will permit him to quickly isolate the most effective mode

of failure prevention.

ACKNOWLEDGMENTS

The author would like to express appreciation for the

support of the Fluid Dynamics Branch of the U. S. Office of

Naval Research during the many years of study at HYDRONAUTICS, which have provided such major contributions to the

present state of understanding of cavitation erosion, and to

again acknowledge Dr. A. Thiruvengadam, who performed most

of this research.

REFERENCES

1. Thiruvengadam, A., "Cavitation Erosion", Applied Mechanics Reviews, pp. 245-253, ASME, New York, March 1971.

2. Thiruvengadam, A., "Prevention and Cure of Cavitation

Erosion", Naval Research Reviews, Vol. XXV, No. 5* pp. 10-

27, Office of Naval Research, Arlington,- Va. , May 1972.

3. Heymann, F. J., "Toward Quantitative Prediction of Liquid

Impact Erosion" , Characterization and Determination of

Erosion Resistance , ASTM STP <W, pp. 212-248, Amer. Soc.

for Testing and Materials, 1970.

4. Thiruvengadam, A., "The Concept of Erosion Strength",

Erosion by Cavitation or Impingement , ASTM STP 408,

pp.22-35^ Amer. Soc. for Testing and Materials, 1967.

5. Proceedings of the Symposium on Erosion, Wear and Interfaces with Corrosion, held June 1973 during the Annual

ASTM Meeting, Philadelphia, Pa. (in press: to be published as an ASTM STP).
DISCUSSION

F. J. Heymann, Westinghouse Electric Corporation : In the nomogram,

along the line that defines erosion strength, you have listed materials.

Is the value of erosion strength given to those materials based on comparative erosion tests, or is it based on calculated strain energy?

A. F. Conn : The history of that goes back to a time when the materials

' first being looked at were eroded out to steady state and there were

some very good correlations between steady state erosion and strain

energy. If I am correct, those are based on the steady state erosion

strengths

.

F. J. Heymann : There are two additional comments I wanted to make.

In using this method, which is, I am quick to say, a good practical

method, it should be remembered that the intensity that you are talking about when you are saying 1 watt/meter2 is not an intrinsic,

independently measureable property of the environment. In fact,

erosion rates vary with the 6th power of the velocity in flow cavitation and with about the same power in liquid impingement, whereas the

kinetic energy associated with the environment varies with the second

power. It's clear that that sort of energy exchange is a very complicated one. We must remember that erosion intensity is a value based

on the amount of material eroded. The other point I wanted to make

pertains to the slide you showed of Thiru's tests where you plotted

the reciprocal of erosion rate versus engineering strain energy for

a number of materials and drew a first order linear correlation. If

those same data are plotted on log- log paper, they will give a better

fit to a second order. That is, the erosion resistance is approximately proportional to strain energy squared, and almost every cor- relation that I've seen shows that erosion resistance is approximately

proportional to the square of some kind of a strength property.

Whether you take engineering strain energy, or ultimate strength, or hardness, there always seems to be some general dependence on the

square of this property. One reason that the ultimate resilience

proposed by Hobbs seems to give a linear proportion is because it

includes the square already. It's S_2_ , and the E doesn't vary that much.

2E

A. F. Conn : Thank you Frank, I agree.

h5
F. B. Peterson, Naval Ship Research § Development Center : I have

two comments and questions associated with them. First, there is a kind of classical experiment of rotating a disc at high speeds

in a non-cavitating situation, where at the outer edge of the disc,

material is removed due just to corrosion. Oxygen reaches various

parts of the surface at different rates. With regard to Thiru's

impeller on the pump of his washing machine, isn't it possible

that the pot metal could have been significantly weakened by corrosion,

particularly corrosion of the outer edge, and possibly cavitation

wasn't even involved in the damage process until the surface was

severely corroded? The other thing is that when a sample is vibrated,

it's my understanding from looking at various peoples' results, that

there is a certain, definite pattern of damage on the surface. Some

people have shown photographically that over a period of time the

number of cavities generated on vibratory apparatus changes, and after

the surface is significantly damaged and is no longer a flat smooth

specimen, the number of cavities is less. In extrapolating the work

on the vibratory rig to a flowing situation, let's say as you showed

on the propeller, or on coatings on hydrofoils, isn't there some difficulty when the hydrodynamics in the flow situation are different

with time?

A. F. Conn : Thiru's washing machine was certainly corroded. The

intensities that I estimated are certainly low enough to be down in

the corrosive range, and corrosion had to play a role. With regard

to that geometry and to the outer tips having high velocity and seeing

more oxygen, the pits were distributed rather evenly over the case

as well as over the impeller. I agree that it was probably a factor,

but not quite the same as the disc. With regard to the key question

about using screening tests and comparing them with other flow situations,

it is not known for sure what effects there would be on screening due

to crevices and too little or too much air, etc. and their relationships

to the characteristic flow, incubation, etc. The very same thing is

seen in foil tests, i.e., the very same rate curve, incubation, and

well defined steady state curve. Comparisons between methods should

be made in the steady state condition.

H. Weinhold, TRW: On your graph of erosion rate vs. cavitation

parameter, you showed a maximum value. It seems to me that could

very well be inherent in the test method. Because of the use of a disc, I suppose you ran all your tests at the same tip speed?

1+6
A. F. Conn : No, this was done for various size foils and it was run for quite a spectrum of velocities. It's the same phenomenon

that was observed around a cylinder flow. It's been observed on rotating discs.

H. Weinhold : The definition of the cavitation parameter, using a pump man's terms, would be essentially net positive suction pressure

divided by velocity head. Obviously, if there is a low net positive

suction pressure, a large amount of bubbles will be generated, but

the collapse energy would be low. At the other end of the spectrum,

a point will be reached where no bubbles at all are generated. There

is a potentially high energy, but there are no bubbles to collapse.

Now I can't see where this is applicable to a propeller. It may

not necessarily be applicable for all cases. I can imagine a pump

which is forced to ingest two phase flow, maybe all condensables,

maybe a mixture of condensables and noncondensables, and force the

fluid into a high pressure regime. The results may not be the same.

A. F. Conn : There are arguments about the temperature effects that

cause this same kind of distribution and the effects on cavitation

number. I am certainly not saying that this thing is going to apply

in every single case.

^7
EXAMPLES OF OIL CAVITATION EROSION

IN POSITIVE DISPLACEMENT PUMPS

John A. Halat and G. 0. Ellis

System Science Laboratories

S perry Vickers

Troy, Michigan U808U

Introduc tion

Investigation over the past twenty years has uncovered certain trends as regards cavitation in an aqueous flowing system that are also useful in understanding cavitation in

power hydraulic systems. However, care must be exercised in

attempting to scale water data to hydraulic oils . In the

hydraulics industry, we are dealing with dissolved and entrained air containing vapor rather than pure vapor. Any

drop below saturation pressure, shown in Figure 1, can cause

bubbles. Since pumps by definition create pressure differences, we can expect cavitation to be a major problem.

Examples of Erosion

Among the variety of hydraulic components, this presentation

is concerned principally with one particular hydraulic device, namely, a piston type, positive displacement pump.

Figure 2 is a schematic of the type of unit that is discussed. The important parts are the reciprocating pistons,

which displace the fluid to produce flow under a pressure

differential, the swash plate, which produces the reciprocating motion when the cylinder block is rotated, the slipper shoes that ride on the swash plate and connect the pistons to it, the cylinder block, the valve plate that

separates the high pressure discharge and low pressure inlet,

and the shaft that rotates the cylinder block. Interfaces

that are sites of potential material erosion, such as the

valve plate -cylinder block, and shoe -swash plate are identified and discussed. In addition, sites such as cylinder

block-piston bore are also touched upon. Some examples of

these parts damaged by oil cavitation erosion will be available for your examination.

14-8
Cavitation Mechanisms

An eight minute film showing cavitation phenomena occurring

in a plexiglass modification to a 20 gpm piston pump is to

be shown. Two mechanisms of cavitation phenomena that were

observed with the model are discussed. The first is the bubble formation associated with pressure regions below the

saturation pressure of air in the system. The second mechanism observed is the bubble formation associated with the

discharge jet at the pump metering grooves located at the

inlet and discharge ports. The effects of design changes as

well as material substitutions are reviewed and their practicality as solutions to the oil cavitation erosion problem

are discussed.

Concluding Remarks

Finally, we shall comment on one observation on oil cavitation erosion that was quite different from what we had expected. Perhaps some are aware of it already. We have a

venturi to study the effects of cavitation and one test was

to study the effect of contaminants. To our surprise, we

found that we could go to a lower pressure before bubbles

formed when we added dirt to the oil.

We are in some disagreement as to why this occurs and time

does not permit a full development of one theory, but, according to van der Waal's equation, a high pressure vapor

could exist inside small bubbles in a metastable state. Thus,

a very small decrease in pressure is required to cause the

bubble to expand. Adding the dirt destroys the metastable

condition and a much lower pressure exists inside the bubble

so a much lower pressure outside the bubble is required before the bubble expands.

k9
50

DISCUSSION

F. B. Peterson, Naval Ship Research § Development Center : When I think of van der Waals' forces, I think of the very short range

forces between molecules. They may be dispersion forces, dipole

forces, etc. These forces act over distances of about 3 x 10" 8 cm.

It would seem to me that in the vapor there would have to be a very

high density of molecules in order to have the molecules that close.

In other words, if it's a gas phase, I don't see how the molecules

would be that close because that's on the order of molecular sizes.

G. 0. Ellis : There must be some transition from the liquid to

the vapor. I can't see that liquid instantaneously changes to vapor.

F. B. Peterson : There is an unbalance in the forces between the

molecules. A molecule out in the middle of a fluid, with essentially

no other types of molecules around, is going to have a kind of

balanced force field. But the molecule that's at an interface is

going to have an unbalanced force field, and its attraction is going

to be greater in one direction than in others. I don't see the

problem of having molecules extremely close. There is a constant

interchange, but I don't see why it has to be a smooth transition.

H. Weinhold, TRW: Frankly, I don't think it is a problem for a physicist. It's a simple engineering problem. In one of the pictures

shown, the worst erosion occurred at the bottom dead center. There

was little erosion at the top dead center point, the transition from

the discharge point to the inlet point. In the movie, what you term

as your metering slot is very reminiscent of a diffuser. It is

probably only a moderately efficient diffuser. It appeared that

when bottom dead center was reached, a stream of bubbles began to be

ejected. Now, if those things are put together, it's very reminiscent

of a cavitating venturi-a venturi forced to work in the cavitating

regime. If it was a highly efficient diffuser there could be a high

inlet pressure. The pressure would decrease to the vapor pressure

at the throat and maybe 90% of the velocity would be recovered, so there would be a total pressure drop of 10%. The bubbles collapse

some place in the wheel. I don't know why you have a tapered diffuser

shape featuring a metering slot in there at all. Maybe you need it;

maybe it's just for manufacturing reasons. Tell those guys to jump

in the lake and put a flat slot in there. That immediately gives a maximum opening and thus takes the diffusion capability away. That

will probably reduce the problem greatly.

52
J. A. Halat : We agree. We've tried maybe every possible design

we could think of, e.g., ramps, slots, holes. We do reduce the amount

of erosion damage with certain designs. If the material loss is

reduced with this type of a unit, the overall noise produced

(the airborne noise) is increased. Only half the cylinders are under pressure, the other half are not. There is a pressure

balance to hold. If the balance is not held, there are violent

oscillations.

Anonymous : Are you talking about a sound audible to the human ear?

J. A. Halat : Right.

Anonymous : What is the frequency of it?

J. A. Halat : It's in the significant hearing range — 2,000 cycles.

Although I haven't measured it, people in our lab have. It can

be traced directly to modifing timing grooves to overall noise.

53
MICROSCOPIC INVESTIGATION OF CAVITATION EROSION

DAMAGE IN METALS

J.V. Hackworth and W.F. Adler

Bell Aerospace Division of Textron Buffalo, New York 14225

Cavitation erosion results from repetitive, localized pressure pulses applied to a material surface by the

collapse of cavitation bubbles. Since cavitation erosion is a mechanical process, many investigations have

been undertaken to correlate erosion rates with some mechanical property or properties of materials. The

lack of success of these investigations is understandable because the pressure pulse generated by the collapse

of a cavitation bubble is applied over an area that is microscopic in size (estimated to be on the order of

0.0001 to 0.010 inch in diameter depending on the source of the cavitation). Thus, the bulk mechanical

properties of the material specimen would have little bearing on the material's response to localized pressures applied at the microscopic level. The microscopic properties and microstructure of the material would

be expected to dictate the form of the erosion damage mechanisms and to control erosion rates.

This paper presents the results of research to identify the cavitation erosion damage mechanisms at the

microscopic level for three metals (aluminum, stainless steel, and titanium) representing a range of properties

and microstructure. The metals were exposed to cavitation generated in distilled water by a 20-kHz ultrasonic facility operating at a vibration amplitude of 2 mils. Representative properties of the metals and experimental details are summarized in Tables 1 and 2, respectively. Replicas of the eroded surfaces of the speci- mens obtained periodically during exposure were examined with a transmission electron microscope to

follow progression of the erosion damage and identify dominant erosion mechanisms as a function of exposure time. Eroded surfaces of selected specimens were also examined with a scanning electron microscope to

assist in the interpretation.

1 100 Aluminum. — Scanning electron micrographs in Fig. 1 illustrate the acretion of bubble collapse

pits on the surface of 1 100 aluminum. The pits occurred at random locations in the initial stage of the

erosion process (0 to 45 sec); however, multiple impacts at previously formed pits became apparent as exposure times exceeded 45 sec. This stage of multiple impact is illustrated by the transmission electron micrographs in Fig. 2. Repeated impacts within preferential pits enlarged and deepened these initial pits. The

walls between these relatively large pits became thinned and were extruded upward from the surface by 120

sec. Fracture of these thin walls between pits became apparent after exposure of 150 sec. as illustrated by

the encircled region in Fig. 3. The surfaces of these fractures exhibited dimples characteristic of ductile

rupture as shown by Fig. 4.
316 Stainless Steel. — Progress of cavitation erosion of 316 stainless steel in the initial stage is illustrated in Fig. 5. Pressure pulses generated by bubble collapse were generally of insufficient magnitude to

form indentations on the surface as was the case with pure aluminum. The first noticeable changes in surface

topography was localized delineation of grain boundaries (5a) and formation of a few scattered slip lines

(5b). Removal of small chips of material became evident at grain boundaries after 12 min. exposure (5c) and

at slip bands after 30 min. exposure (5d). Erosion by chipping at grain boundaries and slip bands within

grains grew more extensive with time (5e and f). The fracture surfaces showed no evidence of ductility or

fatigue striations. Removal of each chip appeared to occur as a single event associated with a single bubble

collapse pressure pulse.

Titanium-6A1-4V. — Progress of cavitation erosion of Ti-6A1-4V is illustrated in Fig. 6. No evidence of

surface impact pitting was found. The first observable erosion damage was cracking and chipping at phase

boundaries and inclusions (6a). The chipping increased with exposure time (6b and 6c) and became extensive by 48 min. (6d). The nature of the erosion after 48 min. is readily apparent in the scanning electron

micrographin Fig. 7. The fracture surface features, more apparent in the transmission electron micrograph

(6d), show that the small chips were removed by brittle fracture. The removal of each chip appeared to

result from a single impact associated with collapse of a single bubble. Fig. 7a also contains a relatively large

and deep pit of the type which first appeared at an exposure of 36 min. and became more prevalent as exposure increased to 120 min. The striations in the bottom of these pits as shown in Fig. 8 indicate fatigue

fracture.

Erosion Damage Mechanisms. — Results from the microscopic analyses provide the basis for defining

the primary cavitation erosion mechanisms operative in the three metals investigated. Bubble collapse pressure pulses readily deform the low-strength 1 100 aluminum forming pits which became preferential sites for

subsequent bubble collapse. The walls between these preferential pits grow thinner and are extruded outward forming lips as the pits enlarge and deepen during this incubation period. Shearing of the pit lips by

subsequent pressure pulses appears to be the erosion mechanism for pure aluminum.

The magnitude of most of the pressure pulses is insufficient to indent the surface of the 316 stainless

steel and impact pits, per se, do not contribute to the erosion of the stainless steel. However, the magnitude

is sufficient to induce shearing in favorably oriented grains. The incubation period consists of the relative

displacement of grains across grain boundaries and the formation of slip bands within the grains. These grain

boundary steps and slip bands concentrate the stresses produced by the pressure pulses so that localized

fracture of the material occurs with subsequent removal of small particles which serve as nucleation sites for

the development of larger pits. The radial growth of these pits by fracture of small particles at the outer

periphery appears to be the erosion mechanism for annealed 316 stainless steel.

55
The sources of stress concentration in the Ti-6A1-4V alloy are the preexisting interfaces between the

different phases within the grains. Cracks which develop at these interfaces during incubation serve as stress

concentrators for the pressure pulses to initiate material removal by localized fracture. This initial erosion

mechanism appears to be similar to the mechanism for 316 stainless steel. However, a second mechanism

based on fatigue fracture becomes operative as exposure time increases. The occurrence of dual erosion

mechanisms in the titanium alloy illustrates the complexity of cavitation erosion.

This investigation demonstrates a rational approach for isolating the material characteristics wh'ch

most significantly affect the material's response to cavitation. The results show that response of a material

exposed to cavitation is not governed by mechanical properties evaluated on a gross scale, but by microstructural features which interact with the erosive environment. Definition and understanding of the erosion

mechanisms and the controlling microstructural features provide the base required for the development of

erosion-resistant materials.TABLE 1. PROPERTIES OF MATERIALS INVESTIGATED

Material Condition

Ultimate

Strength

(psi)

Yield

Strength

(psi)

Cavitation

Erosion

Rate*

Mils/Hr

1100

Aluminum Annealed 13,000 5,000 1.3

316

Stainless Steel Annealed 85,000 35,000 0.5

Ti-6A1-4V Annealed 136,000 131,000 0.2

*"Steady-state" erosion rate in distilled water in 20-kHz ultrasonic facility.

TABLE 2. EXPERIMENTAL DETAILS

Mean Depth

Specimen of Penetration

Location Maximum at Max.

During Exposure Exposure

Material Exposure Time (min.) (Mils)

1100 Mounted 1 mm 3 0.14

Aluminum under horn tip

316 Mounted on 90 0.20

Stainless Steel horn tip

Ti-6A1-4V Mounted on 120 0.13 horn tip

56

Figure 3. Fracture of Pit Lip

in 1 1 00 Al

Figure 4. Pit Lip Fracture Surface

d. 30 min e. 42 min. f. 60 min.

Figure 5. Cavitation Erosion of 316 Stainless Steel
Figure 6. Cavitation Erosion of Titanium - 6A1-4V (5000X)

Figure 8. Fatigue Fracture Apparent after 36 min. Exposure

59
DISCUSSION

G. C. Gould, General Electric Company : I would first like to compliment you on doing an incredibly nice job. I saw fatigue striations

one time and I believe that the alloy was again a titanium alloy.

Do you know which phase, or is there a single phase, that seems to

be the one that is breaking out by fatigue? Do these cusps or individual particles that break out have a constant crystallography

among them? That is, are they one single grain or one single portion

of the phase?

J. V. Hackworth : In answer to your first question, I really don't

know. It is my impression from spending hours at the microscope

that one particle is eroding preferentially in one phase, and that

is the next thing we are going to look into. Based on your discussion

this morning, I suspect maybe it was the a phase because the advantage

of the 3 phase is that it has more shear planes.

G. C. Gould : That was just what I was driving at. It would be nice

if you could support that.

J. V. Hackworth : The 3 alloy is the more fabricative. As far as the

pieces coming off, I think they are about the size of the pits in the

aluminum, the micro-jet type size, and we haven't tried to collect

these particles and analyze them.

C. M. Preece, State University of New York, Stony Brook : The first

question I have is in connection with the cracks initiating in the

stainless steel. We found what appeared to be cracks initiating

in aluminum, in copper, in cobalt, and in nickel, and in no case could we find any propagation of them. In other words, it appeared

as though a crack was starting and we thought that with further

exposure we would get a crack propagating. Never did we find this.

They just stayed as a groove, effectively, and sometimes blunted

out and sometimes didn't. I wonder if you found anything like that

in the stainless steel?

J. V. Hackworth : I didn't really find anything in the stainless steel

that I would call cracks. Before I did stainless steel I actually did

the ABS type 4 bronze. There I thought I saw cracks, but as I progressed,

I came to the conclusion that these were slip lines. I think that is

what I found in the stainless steel. Now in titanium, between the phases,

I would use quotation marks around the word cracks.

60
C. M. Preece: But if it's a fatigue crack, it must be a crack which

is propagating.

J. V. Hackworth : Right, but I didn't find a crack like a fatigue

crack. The only striations I found were when large particles had

been removed.

C. M. Preece : In all of the materials we saw, we got the grain boundary

effects that you observed in the stainless steel, and again, this

occured earlier on and they didn't seem to develop into anything. Do

you think that these striations could be stress corrosion type striations

rather than fatigue type striations? These are known to occur in

titanium alloys.

J. V. Hackworth : These don't look like what I have been calling cor- rosion fatigue. What we found for corrosion fatigue in titanium, at

least in salt water, was that the striations aren't apparent. They're

harder to find and they aren't straight and uniformly spaced.

C. M. Preece : Did you find any trace of the phase transformation from

a to 3 phase, or vice versa, as Dr. Gould suggested with some of the

steels and cobalt?

J. V. Hackworth : No, we didn't get into the phase transformation.

C. M. Preece : Did you look for it?

J. V. Hackworth : No. I don't think you would get a phase transformation.

61
LUBRICANT PRESSURE RIPPLING IN DYNAMIC HERTZIAN

CONTACTS INDUCED BY SURFACE ROUGHNESS

T. E. Tallian

J. I. McCool

S CS EP Industries, Inc.

Technology Service Center

1100 First Avenue

King of Prussia, Penna. 19406

Because of surface roughness, the pressure distribution

in a lubricated Hertzian contact will be subject to random

fluctuations in time even though there may be no direct

contact of the rough surfaces. The magnitude of these pressure

fluctuations as measured by their RMS value for the contact

of two-dimensional (long) cylinders is found to vary with

1) the location in the contact 2) the spectrum or frequency

content of the surface roughness 3) the mean plateau film

thickness and 4) the RMS value of the surface roughness.

Figure 1 is a sectional view of the contact zone of a

rough cylinder in lubricated contact with a flat plane.

(The lubricated rolling contact of two rough cylinders of

general radii can always be reduced to an equivalent rough

cylinder and smooth plane configuration.)

The contact width 2a is assumed to be equal to the value

62
calculated for smooth surfaces under the same load. The

surfaces move at a velocity U in the positive X direction with

respect to a coordinate system fixed to the contact. The

separation is a function of x and time, t, expressible as,

H = h(x) + r (x-Ut ) (1)

where h(x) is the mean film profile i.e. the film profile

that obtains when the surfaces are smooth and, r(x) for

fixed x is a stationary, ergodic random process having a 2

zero mean and variance a . The roughness is considered to

consist of grooves running transverse to the rolling direction.

The mean film thickness h(x) diminishes rapidly with x

over the inlet region, is very nearly constant over most of

the contact zone and decreases near the outlet to a value

roughly 3/4 the plateau level, to form the outlet constriction. (The precise dimensions of the outlet constriction

depend upon the lubricant and operating variables.)

The pressure distribution p(x) and the film shape H(x)

must jointly satisfy the Reynold's equation of hydrodynamics

and the displacement equations of elasticity theory.

Under the assumption that lubricant viscosity is independent of pressure, the relevant form of Reynold's

equation reduces to,

63
where q ( x ) is the isoviscous pressure distribution.

If viscosity varies exponentially with pressure, as

T]= T)^ exp ( ap )

f the pressure distribution p(x) may be obtained

from the isoviscous solution as,

/ . 1 1 p(xJ = - log /o-i * 1-a q(x) K6i

Eq. (2) may be integrated to give

o S q

IT — = 12 n U(h-h) (4)

dx o

where h is the value of the mean film thickness at which

H

3 15 = 0 .

Following the Grubin assumption, over most of the high

pressure (plateau) region q is constant with value l/<*, and

hence ~ 0 and therefore ^3. Z Q and h = h

f irrespective

dx ~ 3x

of H over this region. We may thus regard the right hand side

of Eq. (4) as non-random and express llH as,

dx

3q -3 — = A(x) . H (5)

dx

where A(x) = 12 11 U(h-h) is a non-random function of x.

— o We now find the variance of p(x) by the following series

6h
of steps:

1. Assume a spectrum for the roughness process r(x).

2. Calculate the auto-correlation function (ACF) of

r(x).

3. Calculate the auto-correlation function for H~° =

-3

(h+r) in terms of the ACF of r(x).

-3

4. Calculate the spectrum n(u)J of H as the Fourier

transform of the auto-correlation function.

5. Calculate the spectrum of dq/dx

as A tH^).

6. Calculate the spectrum of q(x) using the relation

between spectra of functions and their derivatives

7. Calculate the variance of q(x) by integrating its

spectrum over cu.

8. Calculate var p(x) = exp2 ap. var q(x).

In performing the calculation it is assumed that the

s Pectrum of r(x) is bandwidth limited white noise of the

form,

tt (cu) = K for UU -b < m - M

n + b

r

0

0 = 0 elsewhere

wherein K is selected so that the integral of the spectrum

2a • 1/2

The final expression for var p is,

65
var* /^p = 29exp( cm) T] u 1 - ~ • & — • o h

3 h cu 0 f

1 + 20Co/h)2 ] mv^)

2 1 1 I

1/2

Existing numerical solutions of the EHD problem for

cylinders do not give sufficient discrimination to determine

1/2 maximum values of var ' p very precisely. This is evident

in the plot shown in Figure 2 of the quantity ! iLlE! . expap

h

4

to which var^/^ p is proportional, constructed using data

from (1).

Using the values,

I h — hi 0 /0 -jJ expop = 2/3 (inlet)

= 2 (constriction)

1/2 will give under-es timates of maximum values for var ' p

for the inlet and constriction regions.

1/2 Figure 3 shows (3 var p) thus under-estimated for

max

the inlet and constriction region plotted against h/a for

b/a>

0 = 0, 0.8 and 0.95, calculated using U = 460 in/sec,

= 10" 4 in 2/lb. Ti = 14.10"7 lbsec/in 2

f p = 4 x 105 psi,

'o max

h = 6000/in and h = 11 x 10~ 6 |Un.

Ripple amplitude at the constriction ranges from

7.104 psi to 6 x 105 psi at h/a = 2.0.

Ripple pressures calculated using the assumptions of this

66
analysis are thus seen to be of an order comparable to the

maximum Hertz pressure.

REFERENCES

1. Cheng, H. S., "Calculation of El as t ohy drodyn amic Film

Thickness in High Speed Rolling and Sliding Contact",

MTI Report No. 67TR24, 1967.

67
68

h / <or

Figure 3.

70
DISCUSSION

J. A. Walowit, Mechanical Technology : What is going on right in

the center of contact where there is 400,000 psi pressure? It

looked to me like a typical asperity could have a height of the

order of half the thickness of the film and a length of maybe

20% of the half-width of the contact zone. If that is the case,

how would that perturb the pressure right in the contact zone?

How would that affect neglecting the pressure gradient, the

influence of the asperity on the pressure gradient term in the

Reynolds ' equation?

J. I. McCool : We did make the assumption that the roughness

frequencies were large relative to the contact width. I don't

know how you made that calculation.

J. A. Walowit : I was trying to take the number of waves per inch

that you said were present in the example. That's about 1000, or about one mil per wave, with a half-width of 10 mils. That's 10

to 1 right there and it could easily have been 2 mils per wave

which would give 5 to 1. It looked like a significant fraction

just from that.

J. I. McCool : Naturally, the validity will erode down in that

regime. We feel best in saying that these results will apply where

frequencies are high relative to the contact width. This would

get back into studying asperities as individual things, which is some- thing that was done at Mobil a few years ago. It was found that where

two individual asperities collide, there were staggering magnitudes

of temperatures and pressures.

L. L. Ting, Ford Motor Company : In his original model, Christian

considered three types of roughness: widths, transverse and

longitudinal, and so-called isotropic uniform. It seems to me

in your analysis you only solve for the transverse case. Are you

going to extend your analysis to consider surface roughness independent of direction, just like the uniform isotropic case?

71
J. I. McCool : I certainly think it deserves attention because

obviously there will be flow around these things. Maybe we've

over-estimated the actual situation by neglecting this flow.

In three-dimensional contacts there may be a circular contact

with a horseshoe around it. By analogy, this sort of thing

could be expected to occur there, but in that case, neglecting

the crosswise components of flow could lead to erroneous results.

Initial results indicate that the effects can be large. Now

we must go further.

F. G. Hammitt, University of Michigan : With regard to the pressure

coefficients on viscosity, I am running through some rough numbers

from the view point of impact of droplets. In erosion or cavitation

damage, one gets pressures in the range perhaps of 10 4 or 10 5 psi

from straight- forward calculations. If the coefficient that you

showed is applied, there would be an enormous increase of viscosity

which could then affect the behavior of the fluid under those

conditions rather drastically. That was for oil, I suppose. Was

water at all similar in this regard?

J. I. McCool : Definitely in lubricated EHD contacts, the viscosities

do rise astronomically over their ambient values. We consider the

fluid to be more like a tar than a liquid as we know it.

F. G. Hammitt : That does not very much affect the calculations of

pressures under droplet impact or cavitation bubble collapse.

J. I. McCool : I hadn't realized that the EHD and the cavitation

people hadn't converged. Maybe this is something the two disciplines

can get together on, because we have been considering pressure rises

with viscosity for years.

F. G. Hammitt: Do you know if water behaves in a relatively similar

way to oil?

J. A. Walowit : I think it's considerably lower than the oil. Oil can

have a pressure coefficient of viscosity of 10~ k reciprocal psi, which

means that a 10,000 psi pressure rise will up the viscosity by a factor

of e. That means 100,000 is at e

10 and 400,000 - it's ridiculous. Eventualy one of a number of things will happen as these enormous pressures

are approached: 1) the viscosity curve will level off, 2) it's been

conjectured that there will be solidification. Water is a somewhat

different animal in that it has a much lower pressure coefficient

of viscosity.

72
F. G. Hammitt: Do you have any idea what magnitude it is for water?

J. I. McCool : We never use water on purpose.

J. A. Walowit : You're right. Water is something you do not put

in your rig. I'd say it's many orders of magnitude lower, but

I don't know.

F. G. Hammitt : In the same regard, that's a steady state effect,

isn't it?

J. I. McCool : Yes.

F. G. Hammitt.' What about very high rates of erosion? Those also

have the effect of making the viscosity apparently go up. Is that

also true for oil?

J. I. McCool : Yes, there are models for that.

E. E. Klaus, The Pennsylvania State University : There are several

problems in viscosity-pressure. One is the effect of temperature.

The effect of 400 psi pressure on the viscosity is about equivalent

to the effect of one degree temperature change. The pressure rise

causes an increase and the temperature rise causes a decrease. It

is true that at the very high pressures, there are very high viscosities. There are some classic data in the literature on this,

and there is a lot of current activity trying to find out what

happens in transients. The EHD film behaves as though the viscosity

is very high once it gets in the region that Jack McCool was showing

in his slides. It's the inlet zone that people are talking about

as to how much fluid will get into the conjunction. Most of the

treatments, until quite recently, have been isothermal, not accounting

for temperatures. Recently the people at Georgia Tech have tried to

measure temperatures in the kind of contact we're talking about here

and have found temperatures of the order of 600 °F on the surface of

the bearing during this occurrence. It's a little more difficult to

measure the temperature in a fluid, but if it is assumed that the

fluid temperature can also be high under these conditions, the two

effects fight each other. The effect of pressure causes a very sharp

increase ih viscosity and the effect of temperature causes a rather

dramatic decrease.

73

SESSION III

!| Chairman: M. J. Devine

j] Naval Air Development Center

1
I
ASPERITY LUBRICATION AND CAVITATION IN

FACE SEALS

J. A. Walowit

Mechanical Technology Incorporated

Latham, New York 12110

The problem of paralle 1- surf ace lubrication has been of considerable interest to people involved in rotary-shaft face-seal and thrust-bearing

technology for many years. The basic aspect of parallel-surface lubrication that distinguishes it from other areas of lubrication technology is

that classical lubrication theory does not predict the existence of a stable hydrodynamic film for steady- state , isothermal, incompressible flow

between smooth, parallel surfaces. Hydrodynamic films between apparently

parallel surfaces have been observed in practice and are often essential

for the reliable performance of thrust bearings and seals. In order to

account for this fortunate discrepancy between classical theory and ex- periment, various investigators have proposed load- support mechanisims

relating to accidental features characteristic of seal performance, which

relax one or more of the assumptions in the classical theory and permit

the theoretical prediction of load support. Some of the features that

have been analyzed are vibratory effects such as wobble and bounce, surface waviness, nonsymmetric rotation resulting from various types of misalignment, lubricant density change, non-Newtonian lubricant effects, and

surface roughness. A partial bibliography of these analyses is given in

References 1 and 2. Most of the mechanisms dealt with in these analyses

can account for various modes of seal performance when the hypothesized

mechanism is present.

In initial experiments performed in the research described in Reference 1,

efforts were made to construct laboratory face seals having smooth parallel surfaces and to keep the accidental features discussed previously to

a minimum. In keeping with the classical theory, the initial result was

77
that these seals did not develop any hydrodynamic load support. After

the seals were run for several minutes, however, hydrodynamic load support did occur simultaneously with the development of cavitation streamers initiating at the many scratches that were working into the surface

during the initial running period. Cavitation streamers have also been

observed on practical sealing surfaces such as the carbon graphite stator

shown in Figure 1. The carbon graphite stator was lapped flat to within

5 microinches, rms . The dark regions are oil and the light regions are gaseous. The points at which some of the streamers appear to terminate

and originate abruptly correspond to irregularities on the stator surface

indicating that streamers were associated with surface roughness.

Fig. 1 Cavitation Streamers in a Face Seal Containing

A Carbon Graphite Stator and a Transparent

Rotor. Motion is From Left to Right

After a series of experiments performed with smooth and rough lapped copper stators (see Reference 1) it was decided to study the effects of sur- face roughness and cavitation on load support by photoengraving an array

of large regularly shaped asperities having a controlled geometry of cylindrical asperities. A photograph showing both the asperity pattern and

the cavitation pattern about a photoengraved stator is shown in Figure 2.

78
Fig. 2 Asperity Pattern Showing Cavitation Streamers on

a Photoetched Copper Stator. Asperities are

12 Mils in Diameter and 130 Microinches High.

The surface finish on the tops of these asperities were apparently as smooth as the surface that did not support load in the initial experiments with smooth- lapped stators, since the etching process associated

with the photoengraving process attacked the region around the asperities

but not the asperity surfaces. These photoengraved stators were found to

develop stable hydrodynamic films immediately; i.e., no wearing- in was

needed

.

Lubricant film thicknesses over the asperities were determined by measuring the capacitance reactions between the rotor and the stator and were

found to range between 5 and 20 microinches which is small compared with

the asperity heights.

79
The cavitation patterns were relatively steady and the cavities are believed to be filled primarily with gas coming out of solutions as opposed

to vapor. No cavitation damage has ever been observed and the cavities

did not disappear instantaneously upon shutdown.

In general, the pressure profile about an asperity would tend to be positive in regions where the film thickness decreases in the direction of

motion and sub-ambient in areas where the film thickness increases in the

direction motion. These high and low pressure regions should tend to

cancel each other out over an entire surface resulting in little or no

load support were it not for the occurrence of cavitation. Cavitation

will limit the sub-ambient pressure to the cavitation pressure thereby

preventing this complete cancellation.

An analysis of the load support that could be generated over and around

perfectly flat-topped asperities having a geometry such as that shown in

Figure 2 drastically under-predicts measured film thicknesses. Some

rounding has been observed to occur near the edges of the engraved asperities however, and it is possible that convergent portions of the asperity tops could be responsible for the observed load capacity. Cavitation would be expected to occur over divergent portions.

In an attempt to correlate film thickness measurements, the asperity tops were modelled to have a small effective tilt as shown in Figure 3 where

the film thickness was assumed to be small compared with the asperity

height and the tilt 6 = h n - h was assumed to be small compared with

1 o the mean film thickness h . The analysis of this model which is described in Reference 2 resulted in the following formula for load capacity:

_ 3

p = . 75uu R 6/h ^ o o o

where u is the lubricant viscosity, u q

is the linear speed of the rotor

and p is the applied load divided by the area of the tops of the asperities. The effective tilt parameter 6 was determined by a fit to the

80
experimental data to be . 43 microinches which is small compared with

the film thickness.

I

Fig. 3 Theoretical Asperity Geometry

A comparison of the semi-empirical relationship given above and the ex- perimental data is shown in Figure 4. The data were obtained with three

different test machines, two different lubricants [a mineral oil (Sohis

pin 100) having a viscosity of 9 centistokes at 130 F and a. jet-engine

fuel (JP-5) having a viscosity of 1.3 centistokes at 130 F ] five differ

ent stators, and eight different rotor speeds [ varying from 220 rpm

(3.5 fps) to 900 rpm (14.4 fps) J. Although scatter in the data is evident, in general the correlation is good. The standard deviation of the

— 6

film- thickness data from the curve being fitted is 1.25 X 10 in.

81
•.

»'M— .

• • « t

*

• ^.v .

.

•

• *• ; >>1 1

•,v > .

• • •

•

•

^*»^ •

Least-squares curve

(8 » 0.43 microinches)

1 1

I07

io8

(p//iu0),in:'

Fig. 4 Correlation of Film-Thickness Data With

The Parameter ( p/uu ), R =6 Mils.

o o

All of the research summarized here was performed at Battelle Memorial

Institute and is described in detail in References 1-3. It is concluded

that the theory, developed on the assumption that there exist small tilts

on the tops of the asperities, correlates well the experimental data on film thickness for the microasperity patterns studied. Further, from the

load support observed, it is apparent that use of planned microasperities

is an effective method for lubricating the parallel surfaces of face

seals and thrust-bearing surfaces.

REFERENCES

1. Hamilton, D. B., Walowit, J. A., and Allen, C. M- , "A Theory of Lubrication by Microirregularities" , Transactions ASME, Series D, Journal of Basic Engineering, Vol. 88, No. 1 (March, 1966).

92
Anno, J. N. , Walowit, J. A., and Allen, C. M. , "Microasper ity Lubrication", Transactions ASME, Journal of Lubrication Technology, Vol.

90, No. 2, p. 351 (April 1968).

Anno, J. N. , Walowit, J. A., and Allen, C. M. , "Load Support and

Leakage from Microasperity- Lubricated Face Seals", Transactions ASME,

Journal of Lubrication Technology, Vol. 91, No. 3, p. 726 (October

1969)

.

83
DISCUSSION

J. D. Cogdell, Caterpillar Tractor Company : I was intrigued with

your film thickness measurements. I am curious as to how you

got down to such fine measurements with capacitance probes.

J. A. Walowit : We did not use capacitance probes. We measured the

gross capacitance across the film. In other words, the whole machine

the rotor and stator, was the capacitor.

M. Wildmann, Ampex Corporation : If you used the gross capacitance,

how did you take account of the dielectric of the oil and the effect

of the cavity which has a different dielectric from the .oil itself?

J. A. Walowit : This is indeed a problem. We did make the assumption

that the areas over the asperity tops were largely oil, and we did

use the dielectric coefficient corresponding to the oil. The area

around the asperities will not contribute much to the capacitance

because we are talking about 10 yin. over the asperities, and the

asperities themselves are about 100 yin. high. But even over the

asperities there can be cavitation which can lead to some error.

G. 0. Ellis, Vickers Division, Sperry Rand Corporation : Right now

I'm involved with the cameraing on vane pumps and on the vanes

themselves. We have random failures on our production tests and

I am trying to find out why. I've developed a mathematical theory,

much as you have, based on roughness. It was pretty hard to convince

people that some roughness is needed in order to get the vanes to

ride on a film. The vanes have a slanted tip, and the tip has a 0.025 in. radius which will usually generate a film. We studied

surface roughness and we couldn't find much difference between just

a fairly smooth surface and a smooth surface, but we found a lot

of waviness. We first found it in the axial direction with a Talysurf. We obtained a trace of the contour to see if we were

getting the correct contour. We had two machining procedures which

gave us different contours and different waviness. The one with

the larger waviness always failed. They both had about the same mean average roughness, but one had more waviness than the other.

It should be realized that a CLA (center line average) roughness,

doesn't indicate anything about waviness. By waviness I mean that

the roughness is superimposed over a 1/4 inch wave, and it seemed

the wave was the thing that was generating the film.

8U
J. A. Walowit : 1 agree. Surface waviness, if present, is a very

effective film generator, but other factors may have to be considered.

If there is a wedge to begin with, waviness can either be advantageous

or detrimental. It can detract from the wedge action. In my case,

I have perfectly flat surfaces. With a small wave on them they will

support a lot of load. Surface waviness will really work well, and

this is one of the "accidental" features that has been studied quite

a bit in the seal literature. In our particular work on asperity

lubrication, the surface waviness and the asperities will fight each

other. Particularly, the asperity lubrication will provide so many

ambient pressure sources along the wave that the wave itself will

cease to be effective. Furthermore, in our own test rig, we had

looked at surface waviness quite a bit and we had gotton the amplitudes

of the surface waves way, way down.

R. L. Johnson, NASA, Lewis Research Center : I am wondering if, in

your asperity lubrication, you ran under conditions where, by virtue

of cavitation phenomena, you were getting starvation on the asperities.

This could adversely affect the load capacity, the lift obtained from

the fluid at the interface. Another point I want to make is simply

a comment. All of this discussion brings out the very pointed need

for better ways of defining surfaces. Three-dimensional methods of

measuring surfaces are needed to give both texture and roughness, and

also a model for properly describing the surface so that it can be

treated mathematically. This is a very urgent need that I think is

pointed up by work such as this.

J. A. Walowit : In answer to your first question some of those asperities

were certainly at least partially starved. For our own apparatus, though,

we have these tall asperities and a radically interrupted surface.

There are many ways for lubricants to get in, so in some cases there is

probably more flooding than there normally would be. In other cases, because

of the cavitation going right up to another asperity, there is probably

partial starvation. So both are probably entering in.

H. Weinhold, TRW: In one of your schematics you solved the Reynolds'

equation tor a tapered wedge with a circular cross section. How did

you solve the Reynolds' equation? Did you use relaxation techniques?

It sounds like a formidable mathematical undertaking — it must be solved

in three dimensions.

85
J. A. Walowit : A great deal of linearization and simplification occurs based on the assumption that the tilt itself is small compared to the

mean film thickness.

H. Weinhold: So the end leakage doesn't have to be accounted for.

J. A. Walowit : There is a simple analytical formula. It turns out

that the value of the tilt that we used was about 1/2 yin. The mean

film thickness was on the order of 10 or 20 yin. There is a 1/2

to 20 ratio which is indeed a small tilt.

H. Weinhold : So you solved it actually as an infinitely long bearing?

J. A. Walowit : No, we made use of the circular geometry. We could

use essentially a constant H except for the generating term in Reynolds'

equation where we plot the slope.

H. Weinhold : Do you know of any work that is being done to define

surface configurations as produced by various processes? I remember

years ago at an MFPG session there was a gentleman who was doing

some work where he showed mathematically that for the definition of

the surface, the rms, the first derivative, and the second derivative

are needed. Then by statistical analysis he could prove he knew what

the surface was. Has any work been done along those lines? In addition,

has any work been done to non-dimensionalize hydrodynamic effects of

such surfaces? Is there a practical tool available for the designers?

J. A. Walowit : A great deal of work has been done in many of these

areas in heavily loaded contacts as opposed to parallel surfaces, where

again small films are dealt with and the roughness is not necessarily

responsible for the lubrication. It's affecting the process, partially

elastohydrodynamically, bordering between thin film lubrication and

dry contact. There has been a great deal done to classify surface

roughness and to take into account its interactions. There has been

some work done in seal-type geometries in a continuation of the study

that I have spoken about. Special types of controlled shapes of

surface roughness have been looked at.

86
L. L. Ting, Ford Motor Company : Is it necessary to consider elastic

deformation of the micro-asperities in your analysis?

J. A. Walowit : Not these particular ones. We had done that. We

computed the elastic deformation to see if that was interacting with

the tilt. But with a tilt of about 1/2 yin. across the asperity,

to fit the data elastic deformation would give 1/10 to a couple

of hundredths of that magnitude.

87
EFFECT OF CAVITATION ON FLUID STABILITY IN POLYMER THICKENED FLUIDS

AND LUBRICANTS

J. L. Duda and E. E. Klaus

The Pennsylvania State University

University Park, Pennsylvania 16802

Linear type polymeric additives are widely used in modern hydraulic fluids and lubricants to improve the viscosity-temperature

characteristics of the finished product. In some cases these additives

have been developed to provide detergency, pour depression, dispersancy

and lubricity in addition to improved viscosity-temperature properties.

The effects of high energy input from high temperatures, gamma radiation, turbulent flow and ultrasonic irradiation on polymer solutions

are shown as a permanent loss in viscosity of the fluid. In all cases

the initial rate of viscosity loss is high and a stable viscosity above

that of the base stock is approached with time. Both the type of high

energy source and the energy intensity influence the asymptotic

viscosity of polymer solutions. The similarity of time-energy severity

relationships has resulted in a number of studies to determine the similarity or differences in the mechanisms of polymer degradation.

The polymeric additives used in functional fluids includes polybutenes, polyacrylates , polymethacrylates , polystyrenes, polyesters, etc

The comparisons used in this evaluation are based on polybutene and

polymethacrylates. Several molecular weight ranges of polymethacrylate

are represented.

88
A comparison of thermal and mechanical stability tests for these

polymer solutions show distinct differences in the mechanism of the

polymer degradation. The thermal degradation tests appear to cause an

unzippering or depolymerization type of reaction in which there is no

particular selectivity based on molecular weight. The reason for the

selectivity noted by the characteristic asymptotic shape of the viscosity

curve as a function of time is not clear at this point. However, the

mechanical degradation of these fluids appears to be based on molecular

weight. Measurements of molecular weight distributions during mechanical degradation show that for a given level of intensity only molecules

above a certain minimum size are degraded. The independence of the

degradation mechanisms can be shown by conducting mechanical tests onthermally degraded polymers and thermal tests on mechanically degraded

polymers. In each case the second test shows a substantial change in

properties after the initial test had indicated essentially no further

degradation with time.

Radiation is in many ways the most severe of the degradation tests

conducted. This type of stability test still shows an appreciable rate

©f deterioration after extended exposure time. In fact effective nuclear

collisions appear to result in polymer degradation in all cases.

Ultrasonic radiation is probably the most widely used test to

degrade polymer solutions (ASTM Test Method D2603-70) . The mechanism of

polymer degradation in this test appears to be cavitation. The sonic

stability tests and the mechanical stability tests have been compared by

investigators with the general conclusion that there are significant

89
differences between the degradation produced by the two test methods.

In general agreement between sonic and mechanical degradation tests are

relatively good when the same polymer or polymer type is compared.

Comparisons involving polymers of different chemical classes are not in

good agreement. This lack of agreement suggests that the mechanisms

involved in the two processes may be different. On the other hand the

mechanisms may well be similar and the level of severity different.

Several mechanical evaluation tests have been devised to determine the

potential role of cavitation in the mechanical degradation process.

In this study, a sharp edge orifice was employed as a device for

producing mechanical degradation in polymer solutions. The sharp edge

orifice is a simple yet severe device for producing mechanical degradation of polymer solutions. When the pressure downstream of the

orifice is of the order of less than 100 psig degradation is a function

of pressure drop across the orifice and gives maximum severity levels.

If, however, two orifices are used in series or if a capillary is used

to produce a high pressure downstream of the orifice, the severity of

the upstream orifice for mechanical breakdown is greatly reduced. Thes

effects are shown in Table 1.

If one defines the conventional cavitation number as

P-P

1/2 plT

where P = absolute static pressure

P^ = vapor pressure of fluid

p = mass density

90
U = characteristic velocity

then it follows that the degree of cavitation or the tendency to cavitate increases as this parameter is decreased. An increase in pressure

downstream of the orifice will decrease the cavitation number and the

cavitation intensity but will not influence the shear field in the

orifice. The influence of the absolute pressure on the degree of degradation clearly indicates that the degradation is a cavitation related

phenomena

.

Another series of experiments was conducted to demonstrate the

mechanism of mechanical breakdown in an orifice. In this series of

tests a single pass through an orifice is used instead of continued

recirculation through the orifice. In simple one pass tests through the

orifice, depending on pressure and orifice size from 3 to 7 percent of

the maximum breakdown after 5000 cycles through the orifice can be

achieved. By using a direct gas drive on the fluid reservoir feeding

the orifice it was possible to force the fluid to pass through the

orifice close to the edge of the orifice by replacing the fluid in the

center of the orifice with the nitrogen gas. As shown in Table 2, the

use of the direct gas drive gave as much as a two fold increase in

severity of the test orifice. These results show that the critical

position for the polymer degradation is the concentric layers of fluid

flowing by the surface of the orifice. The fluid in the center of the

orifice does not appear to be degraded.

One additional series of single pass tests through the orifice was

conducted to test this mechanism. In these tests the fluid was satu91
rated with, carbon dioxide prior to passing through the orifice. The

test results presented in Table 3 show a 30 to 70 percent decrease in

degradation due to the presence of the dissolved carbon dioxide.

The combined data from the orifice degradation test can be taken

as proof of cavitation as a major contributing factor to the polymer

degradation in this test system. The overall data show that the outer

edge of fluid flow through the orifice is the part of the fluid degraded

on each pass. This region of fluid flow would encounter the maximum

velocity and pressure change on going through the orifice. That is, this

portion of the fluid would be in the cavitation zone. The application

of back pressure on the downstream side of the orifice reduces cavitation tendencies and also reduces fluid degradation. The use of

dissolved carbon dioxide in the system tends to make cavitation easier

to achieve but affects the collapse rate of the cavities thereby reducing the cavitation damage. These data all point to cavitation as the

major degrading factor in the orifice system. The level of severity may

therefore, be the major difference between the ultrasonic and orifice

devices

.

Some additional tests have been conducted to evaluate a more

severe mechanical device than the orifice. The device in this case is

two tapered roller bearings loaded against each other by a series of

springs. In the orifice pressure drops up to 3000 psig were used. In

the bearing the pressure in the concentrated contact at the race and the

roller ranges from about 50,000 to 200,000 psig. In addition, the fluid

films (elastohydrodynamic) in the conjunction are estimated to be thin

92
enough that all of the polymer molecules in this film will be degraded.

Some test results with this bearing rig are shown in Table 2. The

results indicate an increased severity anticipated by the increase in

pressure drop and a change in the relative stability of polybutenes and

methacrylates over that predicted by the orifice data.

93
Bibliography

1. Jolie, R. M. , "Structural Effects on the Viscosity Behavior of

Polymer Solutions", Ph.D. Thesis, The Pennsylvania State University

(1957).

2. Klaus, E. E. and Fenske, M. R. , "Some Viscosity-Shear Characteristics of Lubricants", Lubrication Engineering, Vol. 11 (March-April,

1955) .

3. Klaus, E. E., Tewksbury, E. J., Jolie, R. M. , Lloyd, W. A., and

Manning, R. E. , "Effect of Some High-Energy Sources on PolymerThickened Lubricants" , ASTM Special Technical Publication No. 382

(1964) .

4. LeMar , R. L. and Bootzin, D. , " Discrepancies Between Mechanical

and Sonic Shear Methods in Evaluating Shear Stability of Hydraulic

Fluids", ASTM Special Technical Publication No. 382 (1964).

5. LeMar, R. L., "Factors Affecting Sonic Degradation of Polymer

Solutions", I. & E. C. Product Research and Development, Vol. 6,

No. 1 (March, 1967).

6. Groverman, H. V., "Mechanical Shear Stability of Dilute Polymer

Solutions" , M.S. Thesis, The Pennsylvania State University,

University Park, Pa. (1971).

7. Boyer, R. F., " Styrene Polymers-Degradation", p. 240 in H. F. Mark,

N. G. Gaylord, and N. M. Bikales, eds., "Encyclopedia of Polymer

Science and Technology", Vol. 13, Interscience Div. , John Wiley and

Sons, N. Y. (1970)

.

9h
Eisenberg, P. and Tulin, M. P. /Cavitation" , Section 12 in

V. L. Strieter, ed., "Handbook of Fluid Dynamics", McGraw Hill,

N. Y. (1971).

95
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96
Table 2

ONE PASS ORIFICE SYSTEM FOR MECHANICAL DEGRADATION

OF POLYMER SOLUTIONS

0.016" Diameter Orifice - 2500 psig AP

4% n-octyl polymethacrylate 160 in white oil

Values given are average of 20 tests under each condition

Viscosity in cs. at 100°F. % Vise. Decrease

Test Method Orig. After 1 Pass Due to the Polymer

Direct Gas Drive 30.40 27.26 14.6

Pistons in Both

Reservoirs 30.40 28.0 11.2

Table 3

MECHANICAL DEGRADATION OF POLYMER SOLUTIONS

One Pass Orifice System

n-Octyl polymethacrylate - 160 in white oil

0.016" Diameter Orifice - 2500 psig AP

Percent Gas in Sol. Viscosity cs. at 100°F. % Vise. Decrease

Polymer Type PPM(Wt) Orig. After 1 Pass Due to Polymer

2.5 air 100 19.00 18.15 8.5

2.5 co

2

1207 19.00 18.78 2.2

5.5 air 100 43.32 38.46 14.1

5.5 CO 1090 43.32 39.81 10.4

97
DISCUSSION

R. Y. Ting, Naval Research Laboratory : I have several questions.

One, how do you measure the viscosity of these materials? Two,

what is the molecular weight range you are working with, and for

these molecular weight ranges, is the viscosity still Newtonian

or shear dependent?

E . E . Klaus : These are non-Newtonian fluids. The viscosity is

shear dependent, but that's not what I am talking about. That

shear dependence is a temporary viscosity loss which occurs when

the fluid is under shear. I am talking about low shear measurements.

I can get the same order of viscosity decrease in flowing through

a capillary system at high shear, but when it comes out the other

end, it has the same viscosity it had when it went in. What I am

talking about is putting it through a meat chopper, so that when

it comes out the other end, it has that kind of a loss, and there

is no way to get it back. This is low shear viscosity. As for the

question on the molecular weight range, we are relatively naive on

this and I would say that the molecular weight range is a couple

hundred thousand to a million by the usual methods of measuring

molecular weight. We used to think it was much lower than that

until we added Larry Duda to our staff who set us straight.

R. Y. Ting : The reason I asked about shear dependence is because

when a fluid is sheared or degraded, presumably the molecular weight

distribution changes, especially if the fluid is in the million

molecular weight range. There is actually a shift to a lower mean

molecular weight. When the molecular weight shifts tremendously

because of degradation, then the shear dependence also changes.

Therefore, even though the same thing is measured at lower shear

rates, the transition behavior may be different.

E. E. Klaus : The kind of things that we are talking about are Newtonian up to shear rates way beyond the shear rate used to

measure viscosity. We're not dealing in these extremely high

molecular weights that I assume you are familiar with in your work.

We're dealing with things that are considered relatively stable

as far as shear is concerned. I would say the typical shear required

to show the shear sensitivity in a simple viscosity determination is

something on the order of several thousand reciprocal seconds. Below

that it's Newtonian.

98
G. 0. Ellis, Vickers Division, Sperry Rand Corporation: Some years

ago, we were interested in the cavitation noise in valves. This

noise degrades the oil and damages the valves. To study the valves

I started looking at diffuser curves. As the diffuser angle is

increased, the recovery coefficient goes down, of course, until a

90° half angle or an instant opening is reached. We had a recovery

factor of 1/3. We hung a valve on springs and shifted it to different

positions and in each position we varied the back pressure. We found

that at 1/3 back pressure the cavitation ceased. We had an accelerometer mounted on the valve and measured the frequencies. We assumed

these frequencies indicated the beginning of cavitation. A 1/3 back

pressure on an orifice should avoid cavitation. So if there is a series of orifices, there will be no cavitation.

E. E. Klaus : I think the case where we had the lowest permanent or mechanical shear in the loading with a low shear capillary was at

about that ratio of pressures. Incidentally, I would point out with

respect to the paper on cavitation damage on the valve plate, a long

time ago, more like 25 years ago, we used these polymer fluids as sort of an integrated device to tell how mechanically efficient your

pumps were. The amount of degradation in the fluid from Vickers

pumps agreed quite well with the efficiency of the pump, or the level

of noise of the pump. The lower the noise, the lower the breakdown

and the higher the efficiency, which is another thing that can be

done with these fluids, if you believe it's cavitation.

R. S. Miller, Office of Naval Research: How long do the polymers

in our motor oils last?

E. E. Klaus : That depends on what motor oil you buy and how much

you are willing to pay for it. Polymers can be made that will last

from one oil change to another or they can be made to last until you

drive 100 miles, and the difference between the two is the price.

The higher the molecular weight, the less polymer is used and therefore

the cheaper it is. I don't know what the automotive companies are

doing, but in the hydraulic business, the automatic transmission business,

and other areas, the polymer molecular weights are closely monitored.

It is possible to control the molecular weight of the polymer so that,

in aircraft hydraulics, a 15% viscosity decrease maximum is obtained

from the shearing effect. The polymer molecular weight is tailored

to the severity of the device so that it is always between 85-100%

of the polymer effectiveness. This is not as well controlled in

the commercial market.

99
EFFECTS OF VISCOELASTICITY ON CAVITATION

IN DRAG REDUCING FLUIDS

Robert Y. Ting

Naval Research LaboratoryWashington, D. C. 20375

INTRODUCTION

The flow properties of very dilute solutions of high molecular weight

polymers have attracted considerable attention in recent years. A large

part of this interest involved the drag reducing characteristics of these

dilute solutions in turbulent flows,, A reduction of 70$ in the pure water

friction in turbulent pipe flow can easily be obtained by the addition of

a few parts per million of a high molecular weight linear polymer such

as polyethylene oxide. Such a large effect on the frictional drag is a clear manifestation of the influence of these materials on the overall

flow regime. The details of the mechanisms involved in turbulent drag

reduction are not yet clear, but it is generally believed that the viscoelasticity of the polymer solution plays an important role (l). In view

of the demonstrated effect, a study was undertaken to determine whether

these same solutions would influence cavitation and thus reduce the

material damage accompanying cavitation. Specifically, this report presents the results of an experimental study of the viscoelastic effect of

polymers on cavitation inception. The implication of polymer viscoelasticity on the asymmetric collapse of cavitation bubbles in the proximity

of the wall is also discussed.

EXPERIMENTAL

To study cavitation inception in polymer solutions, a blow-down water

tunnel with short running times was used. Tests were made using l/h and

1/2 inch diameter models of hemispherical-nose cylinders. To accurately

detect the inception of cavitation, a reliable technique was developed

using a continuously operating He-Ne gas laser. The laser beam was adjusted to grazing incidence with the model at the minimum pressure point

where cavitation inception was to be expected. A sensitive photocell was placed at ninety degrees to detect the beam. As incipient cavitation

occurred, the bubbles caused scattering of the laser beam which was picked

up by the photocell. Static pressure near the model in the working sec- tion of the tunnel was measured using a solid-state pressure pick-up.

The signals from the photocell and the pressure pick-up were recorded on an oscillograph. Velocity field visualization was achieved using one microsecond duration light pulses scattered by small polystyrene latex

spheres in the flow. Dark field multiexposure photographs taken at

ninety degrees to the direction of the light beam provided reliable data

100
on the flow velocity. Data are presented in terms of the "incipient

cavitation parameter", This parameter designates the state of liquid

flow in which cavitation appears , and is defined as the difference between

the local static pressure and the liquid vapor pressure, Ps -Pv, normalized

by 1/2 pu^ with u being the free-system velocity. The drag reducing polymers tested included Polyox-WSR-301 and guar gum (J2FP).

RESULTS: CAVITATION INCEPTION

Fig. 1 shows the variation of 0^ with respect to Reynolds number.

Data for tap water indicate that the incipient cavitation parameter first

decreases with increasing Reynolds number until a critical value of about

L. 5 x 105 is reached. Beyond this point Oj_ increases with increasing

Reynolds number. The present data for water generally agree very well

with the previous studies (2,3), and thus lend confidence in the measurement technique developed here. It is also clear from Fig. 1 that o\j_ is

much lower in polymer solutions. This reduction in the value of a-j_, by

as much as 60$, shows a definite inhibition of the cavitation inception

of a submerged body in flows of dilute polymer solutions. According

to the present data, it is believed that this incipient flow-cavitation

suppression by polymers is a real effect of considerable magnitude,

effective over a large range of flow velocities and object sizes.

Fig. 2 shows the variation of 0^ for different concentrations of

Polyox-WSR-301 at Reynolds number 85*000. It indicates a steady in- crease of cavitation inhibition with increasing polymer concentration

until about 50 PPMW (parts per million by weight). Also included in

Figo 2 are data obtained by Hoyt (h). In this case, a similar, definite

inhibition of the inception of jet cavitation was demonstrated in turbulent jetSo

If the percent reduction of incipient cavitation parameter is defined

as oW

- R

I =

1

w

1 x 100$ (1)

a.1

t \W / \P where ( ) are values in water and ( ) in polymer solutions, then the

data of Fig. 2 may be presented as shown in Fig. 3« It is clear that

initially the effect of cavitation suppression increases rapidly with

increasing concentration. Then the effect becomes saturated as Rj

approaches a certain limiting value at higher concentrations.

The concentration dependence of Rj can be well-described by an

equation of the form

RT

I c R

I-m [c]I

+ c (2)

where c is concentration, [c] the intrinsic concentration for incipient

101
cavitation suppression, and Rj_m

the maximum value of Rj. If the limit

of Rj/c at zero concentration is examined, it can be seen that the parameters Rj_m

and [c]j are useful measures of the "effectiveness" of the

polymer additive on a unit concentration basis at infinite dilution. The

validity of Eq. (2) may be tested easily, since a linear relation should

exist between c/Rj and c if this relation is applicable. The data of jet

cavitation shown in Fig. h serve a good example of this linearity.

ELONGATIONAL VISCOSITY AND ASYMMETRIC BUBBLE COLLAPSE

One distinct feature of polymer viscoelasticity is the behavior of

polymer solutions in flows with longitudinal gradients. Consider the

case of a column of viscoelastic liquid stretched with its axis oriented

in the direction of stretching. Since the velocity field in this case

is irrotational, additional normal stresses, existing as tension in the

axial direction and compression in the transverse direction, could be

very high. These stress components represent a high resistance to the

flow. To indicate these high stress-levels, an "elongational viscosity",

T], is usually used (5)> defined as the sum of these stress components

normalized by where T)Qis the zero-shear viscosity of polymer solutions, and T is a constant stretching rate. Fig. 5 shows Tf as a function

of the dimensionless stretching rate X^T, where Xj_ is the terminal relaxation time of polymer solutions. It can be seen that as flow time

increases, a very high resistance may build up to retard this stretching

even at fairly low stretching rates.

This phenomenon, characteristic of high molecular weight polymer

solutions, may have strong implications on material damage effects

caused by the presence of small, transient cavitation bubbles near solid

surfaces. Benjamin and Ellis (6) asserted that cavitation damage is

mechanical in nature and that study of individual collapsing bubbles is

the key to understanding the damage process. One important mechanism

involved in cavitation damage is the action of liquid jets formed on bubbles near the solid surface. The formation and damaging potential

of these jets have been experimentally shown in both a static (7) and

flowing fluid system (8). Recent analysis of asymmetric bubble collapse near a solid surface (9) showed that the jet formed near the end of the

collapse process impinges against the solid surface, and generates

pressure pulses with pressures of the order of 1,000 atm. These pressure

pulses are generally believed to be the source of cavitation damage.

In a liquid jet, fluid elements deform along the axis of the jet and

the flow is of a stretching nature. In Ref. 9 it was shown that during

the later stage of asymmetric bubble collapse the jet velocity may be as high as 100 m/sec. Furthermore, an increment of 35 m/sec in jet velocity

may take place in 0. 02 second. For a bubble with an initial radius of

0.1 cm, the stretching rate may be estimated as X^ ~ 35*000 sec"-'- during this period of t = 0. 02 sec0 For polymers such as polyethylene

oxide with molecular weight higher than one million, the typical terminal

relaxation time, Xj_, is of the order of milliseconds. Therefore, the

102
jet formation process in cavitation collapse near a solid surface repre

sents a case where ~ 30 and t/M. ~ ^* ^n vi ew °f the result shown

in Fig. 5, an astronomical increase in elongational viscosity may be

anticipated and the formation of liquid jets could "be greatly retarded

or inhibited. While experimental confirmation is needed, this analysis

suggests that cavitation damage could be greatly reduced by using poly- mer additives with appropriate viscoelastic properties.

REFERENCES

1. Lumley, J. L. , Ann. Rev. Fluid Mech. 1, 367 (1969).

2. Rouse, H. and McNown, J. S. , Bulletin 32, No. k20, State Univ. Iowa

Ames, Iowa (19^8).

3. Kerman, R. ¥. , McGraw, J. T. and Parkin, B. R. , Trans. ASME 77, 533

(1955).

h. Hoyt, J. W. , presented at the l6th Amer. Towing Tank Conf. , Sao

Paulo, Brazil, August 1971*

5. Ting, R. Y. , J. App]_ Polym. Sci. l6, 3169 (1972).

6. Benjamin, T. B. and Ellis, A. T. , Philo. Trans. Roy. Soc. London, A260, 221 (1966).

7. Naude, C. F. and Ellis, A. T. , Trans. ASME 83, $+8 (1961).

8. KLing, C. L. and Hammitt, F. G. , ASME J. Basic Eng. 3D, 825 (1972).

9. Pies set, M. S. and Chapman, R. B. , J. Fluid Mech. ¥7, 283 (1971).

0.9 —

LU

<t

<J

CL.

0.7 -

p 0.5

0.3

a.5 0.1

O WATER

• GUAR GUM 300 PPMW

x P0LY0X-WSR-30I 50 PPMW

O

0.5

Po

x

o

1.5

O

G

QD

• • •

2.0 2.5

O

3.0

REYNOLDS NUMBER x 10 -5

Figure 1: Incipient cavitation parameter vs. Reynolds number.

103
0.7

10 20 30 40 50 60 70 80 90 100 P0LY0X CONCENTRATION ( PPMW )

Figure 2: Incipient Cavitation Parameter vs. Polymer Concentration

6 8 10 CONCENTRATION (PPMW)

Figure 3: Percent Reduction of Incipient Cavitation

Parameter vs. Polymer Concentration

10k
• CD>

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0\ , <r

105
DISCUSSION

F. G. Hammitt, University of Michigan : In the plot where you showed

time and velocity of the microjet, time was plotted in seconds.

Could this be true? You had a full second required for that bubble

to collapse which doesn't seem reasonable to me.

R. Y. Ting : This is Plesset's data from an article which appeared

in the Journal of Fluid Mechanics. I just took the data and plotted it.

F. G. Hammitt : Usually a few milliseconds is sufficient and the end

of the thing is * matter of a microsecond or so. I thought maybe there was a mistake in the scale, like a factor of 10 3

.

R. Y. Ting : In Plesset's calculations, he ignored viscosity effects,

compressibility, etc. That could be the reason why the calculation

appears as it does.

106
PROGRESS AND PROBLEMS IN EROSION PREDICTION

Frank J. Heymann

Technology Development Dept.

Steam Turbine Division

Westinghouse Electric Corporation

Lester, Pa. 19113

In order to predict the adequacy of a component to resist

the loads put upon it, one must be able to express quantitatively

1) the character and severity of the loading,

2) relevant material properties measured in standardized

tests, and

3) failure theories which relate the behavior measured in

simple tests to the behavior expected under different

or more complex loading conditions.

This paper discusses the problem of erosion prediction

from the above-mentioned viewpoint. Its objective is to

evaluate the present status in each of these three areas

(although they cannot be completely separated) , and to point

out some of the unresolved questions and needs for further

research. It concludes with an estimate of the accuracy

which can reasonably be expected for erosion rate predictions.

The emphasis is on liquid impingement as well as cavitation

erosion.

Characteristics of Environment

The preponderant opinion today is that erosion results

from repeated localized impulsive loadings, due to impacts

of droplets, or, in cavitation, due to impacts of microjets

or of shock waves caused by bubble collapse. The most fundamental way of characterizing this environment would be to

describe the actual stress regime to which the component is

subj ected.

107
The repetitive nature of loading leads to obvious analogies with fatigue } although one should not assume that the

removal of material always occurs by fatigue failures. Experiments have suggested that the mechanism can range all the

way from true fatigue, through repeated plastic deformation

and eventual brittle fracture, to ductile necking and tearing.

The stresses and deformations in the material depend on

the time-varying surface pressures and shear flows associated

with each impact. For cavitation, only the roughest estimates are available. Even for impacts of liquid drops on a

plane surface, the magnitude and distribution of impact

pressures is still a subject of controversy. But even that

is not the whole story, because the deformation and erosion

of the surface will itself modify the fluid forces it experiences. On a deeply pitted surface the impact pressures,

flow patterns, and stresses for any given impact will depend

on where it happens to occur.

Thus , characterization of the environment in terms of the

actual microscopic stress loadings is a very difficult prospect. Even so, there have been some recent efforts to

predict erosion on the basis of simplified models of these

fundamental processes (1, 2), with instructive results.

Lacking exact information on the fundamental fluid-solid

interactions , a more common approach has been to characterize

the severity of the environment in terms of directly measurable fluid-dynamic characteristics. For liquid impingement

it is possible, at least in principle, to measure or predict

the number, size, velocity and impact angle of impacting

droplets. Characterization of such an environment then

depends on developing empirical relationships between erosion

rates and these various parameters. To a first approximation,

the severity of impingement can be represented by the volume

flux of impinging water times the fifth power of the normal

component of impact velocity (3).

For cavitation this approach is more difficult. The number, size and collapse velocities of cavitation bubbles is

not generally determinable. In the vibratory cavitation test,

frequency and amplitude are measured, but correlations based

on them are not transferable to field problems involving flow

108
cavitation. Flow cavitation environments can be characterized by fluid velocity, cavitation number, etc., but the

severity of erosion will also depend strongly on the gross

geometry and flow patterns.

The difficulty of independently determining the severity

of a cavitation environment has led Thiruvengadam to characterize it in terms of the erosion it produces (4, 5). This

approach is of practical value when correcting a field

problem or when model tests of a projected system can be

carried out. Thiruvengadam 1 s method is based on an energytransfer concept and yields a cavitation intensity (analogous

to acoustic intensity) having units of power/unit area.

While valid within the context of his method, at present this

parameter is not susceptible to independent measurement.

Characteristics of the Material

Because of the relative inconvenience of conducting

erosion tests —especially of long duration—investigators

have long sought to correlate erosion resistance with tensile

test data, hardness, or other independent material properties

which the designer can more easily determine. But all such

correlations, when tested against a wide spectrum of materials, show scatter ratios of 10:1 or more (3, 6). We must

accept that erosion resistance is an independent property and

must be measured in erosion tests. There is now an ASTM

Standard Vibratory Cavitation Erosion Test (ASTM G32-72), and

a current round-robin test program of liquid impingement facilities is expected to lead to recommended practices and

reporting procedures, if not a standardized apparatus.

The raw result of an erosion test is a curve of material

loss versus time. Its typical shape is now familiar (7): It

begins with an "incubation period" with no material loss but

increasing plastic deformation of the surface. This is

followed by a rise of the erosion rate to a maximum, during

which erosion pits proliferate. From the maximum the erosion

rate gradually falls off (possibly leading to a terminal

steady-state value) , during which the number of pits remains

fairly constant but they enlarge (8) , thereby attenuating the

severity of the interactions.

109
In order to merely rank materials, many investigators have

characterized the result by one number- -most often the maximum erosion rate. It is obvious, however, that if we want to

predict erosion-time histories in the field, we need a mathematical representation for the complete curve which, of

course, must involve several parameters. Depending on the

completeness and on the approach, from 3 to 5 parameters have

been required to formulate long-term erosion-time curves

(3, 9, 5, 10).

In all cases, some of these parameters pertain to the

shape of the curve and some to the actual magnitudes. Now,

the crucial question is this: Are the fundamental proportions of the curve invariant, so that the whole curve can be

predicted by one material property and one environment intensity property? Or are different parameters influenced by

different material characteristics and different environmental characteristics?

At present, there is good reason to believe that the material must be described by at least two properties. For

instance, the results of the recent ASTM Liquid Impingement

Round Robin suggest that the incubation time is affected by

both a different material property and different environmental properties, than the maximum erosion rate.

All three of the above-mentioned models include as one

parameter the mean depth of erosion at or near which the

attenuation of erosion rate begins. An important practical

question is how this depth can be estimated. From some test

results, its value appeared to be roughly independent of the

material, but possibly dependent on drop size in liquid impingement (3, 5). However, comprehensive or conclusive data

are still lacking.

Another important question is whether there exists a

threshold erosion property, much like an endurance limit in

fatigue. Field experience as well as some experimental investigations have suggested that there is threshold velocity

for liquid impingement erosion, but several related questions

remain: 1) How is this velocity to be reliably determined?

2) Does it depend only on the material or also on impingement

conditions? And 3) can the material parameter involved be

110
related to those measured in a standard erosion test? Impact stresses, in a first approximation, should be independent of the size of impacting droplets. Yet experimental

data indicate that threshold velocities (as well as erosion

rates) are dependent on drop size. A good deal of additional

experimental work is needed to settle the threshold question.

Erosion Rate Models

In the erosion context, the role of a "failure theory" is

taken by a mathematical model which relates the parameters of

the erosion-time history to the characteristics of the environment. Some aspects of this question have already been

discussed under the previous headings.

It must suffice here to say that for impingement erosion,

there are vast amounts of data of the dependence of maximum

erosion rates and incubation times on impact velocity and

angle; less so on drop size effects; and still less on the

physical properties of the liquid. More of the latter is

needed before we can use test data obtained with water at atmospheric conditions to predict erosion at elevated pressures

and temperatures, or with different fluids.

For cavitation, on the contrary, generalized laws based on

hydrodynamic parameters are more difficult to obtain, but the

convenience of the vibratory cavitation test has led to much

research on the effects of pressure, temperature, and corrosive interactions, with various fluids including lubricants

and liquid metals.

Accuracy of Prediction

In dealing with erosion test results and predictions, we

must learn to accept uncertainties greater than we are normally accustomed to. One reason for this is that the range

of erosion resistance is much greater than that of convention

al strength properties. Ref. (9) shows "normalized erosion

resistance" values, relative to 18-8 stainless steel, from

many test data in the literature. The overall range covers a

ratio of about 2000 to 1, and the scatter band for any one

material covers a ratio of typically 3 or 4 to 1 . This

scatter derives from differences among test materials,

111
differences among reference materials, and differences among

test conditions.

Among the most closely-controlled tests so far have been

the two ASTM "round robins" of vibratory cavitation devices

and liquid impingement devices, where samples from the same

bars, for each of several materials, were tested under relatively similar conditions. The normalized erosion resistance

values from these tests exhibited "scatter ratios" of 1.5 to

3. It can be shown that, on a logarithmic scale, this

scatter band is about the same proportion of the overall

range of values as is found, for example, for tensile

strengths or for finite fatigue life-times.

Consider, furthermore, that erosion rates typically vary

with the 5th power of velocity in liquid impingement as well

as flow cavitation. If erosion rates are so sensitive to

velocity, it is to be expected that they are equally sensitive to many other factors which affect the fluid-solid inter

actions or the material's response to them. In the ASTM

impingement round robin, erosion rates for the same material,

at the same impact velocity when normalized to equal quantities of impinging water, still exhibited scatter ratios of

about 5:1 among different facilities.

All of these considerations, taken together, suggest that

the best accuracy which we can expect at a high confidence

level, even from much improved prediction schemes, is within

a factor of 2 or 3.

1. C-H Mok, AIAA J

.

, Vol. 7, p. 751 (1969).

2. G. S. Springer and C. B. Baxi, AFML-TR- 72-106 , 1972.

3. F. J. Heymann, ASTM STP 474, 1970, p. 212.

4. A. Thiruvengadam, Cavitation in Fluid Machinery, ASME,

1965, p. 32.

5 e A. Thiruvengadam, Proc . 3rd Conf. Rain Erosion, Royal

Aircraft Estab., 1970, p. 565.

6. F. G. Hammitt et al, ASTM STP 474, 1970, p. 288.

7. F. J. Heymann, ASTM STP 408, 1967, p. 70.

8. Y. Suezawa et al, J. Basic Engineering , ASME Trans

Vol. 94D, p. 521 (1972).

9. J. Tichler and A. W. J. deGee, Proc. 3rd Conf, Rain

Erosion, Royal Aircraft Estab., 1970, p. 847.

10. A. Thiruvengadam et al, ASTM STP 474, 1970, p. 249.

112
DISCUSSION

F. B. Peterson, Naval Ship Research $ Development Center : I would

like to ask a question related to designers choices. If I was responsible for the design selection of a propeller going on a brand new large container ship, I would be faced with the problem

of deciding whether I should invest in a propeller that has a different material than is typically used now from the standpoint

of erosion resistance. I'd have the trade-offs of how often the

propeller might have to be changed and the cost of dockside time

and all the economics associated with the life of the propeller

from the erosion point of view. On the other side of the coin I know that the cavitation performance of the propeller can be changed

strictly by hydrodynamic design, and. that also has its financial

trade-offs. Now, as a designer, if I came to you and told you

what I can do hydrodynamically, and I can lay out all the costs for

you, such as that of changing a propeller, etc., and give you the

flow condition around the propeller, could you tell me what material

I should use and what it would cost or how long the material would

last for a certain extent of erosion?

F. J. Heymann : A quick answer to that is no, I could not. The whole

point of my talk is that there are still an awful lot of uncertainties

that need more work before they can be tied down. I would imagine the

best approach might be doing model tests such as Thiru has suggested

and using scaling approaches and measuring erosion on a model propeller

to predict the prototype performance.

F. B. Peterson : It's really impossible model-wise to scale cavitation

correctly from model to full scale. Model users either lose the dynamics

of the collapse and get the correct pattern of cavitation, or they

get the correct dynamics of the collapse in terms of energy but have

nowhere near the correct cavitation pattern. Therefore, a full scale

prediction based on model work can't be made.

J. F. Corso, Naval Ship Engineering Center : It appears that there

must be some correlation between corrosion resistance and erosion

resistance. We found that alloy 20, while it's not a hard material,

is superior from a corrosion and erosion standpoint to many harder

materials. Could you comment on this and also, is your erosion

prediction analysis valid if you don't take into account debris that

might be in the effluent?

113
F. J. Heymann : Again I would like to say I have not presented an erosion prediction method. I have looked at what is required for

a successful erosion prediction method. I have tried to show that

we are still pretty far away from any such thing. I'm sure that

it is very true that in many instances the corrosive influence is

very significant. It's not always the harder material that is best.

A plot of corrosion test results vs erosion test results taken from

a Russian paper shows a bit of a correlation, but not a very significant one. I'm sure it depends on the particular environment. If

it's one where corrosive influences are important, and the material

is susceptible to corrosion, then that becomes a major factor to be

considered.

A. Thiruvengadam, The Catholic University of America : My comment

is in connection with Dr. Peterson's remark. In this day and age

when we have seen people landing on the moon, it's rather discouraging

to say it's impossible to model erosion in a propeller, etc. This

is one area we have been looking at. We cannot really say we have

simulated all the conditions a prototype propeller would see by

using a model. Given the time and encouragement in terms of research

effort, I believe it could be accomplished.

J. I. McCool, SKF Industries : Why do you classify materials resistance

to erosion as a derivative, a time rate of change, rather than the

time until it achieves a given depth of erosion, the integral of the curve rather than the slope? It seems that would decrease the scatter

and some of the hardships you pointed out with respect to the different

shapes of curves that you got. It also seems that that would be the

thing that a designer would care about — time until depth d.

F. J. Heymann : You've got a good point in the desirability. One thing

that I tried to do in one paper was to base correlation not on instantaneous erosion rates but on average erosion rates and plot them vs. the depth of erosion. Take the slope, which is the average erosion

rate necessary to reach a certain depth, and plot that vs depth. A

time can then be predicted for any number on the curve. But it turns

out that there is a good deal of scatter. We approximate the curve by an incubation time and a maximum erosion rate. It's been my

experience that using this value as a parameter for the erosion resistance of material gives fairly consistant results between different

test apparatus. There is much more scatter in the incubation period

than there is in this. Our ASTM tests showed that. So any sort of

combined thing that gives an integrated value which includes the incubation period will give more scatter. That is one of the difficulties

in predicting long term erosion. I don't know yet what all the sources

of that scatter are.

Ilk
SESSION IV

Chairman: M. L. Picklesimer

National Bureau of Standards

(substituting for W. A. Glaeser)

POTENTIAL OF THIN, SPUTTERED FILMS AS

EROSION RESISTANT PROTECTIVE COATINGS

Konrad Gentner

Naval Air Development Center

Warminster, Pennsylvania 1897^

James E. Newhart

Naval Air Propulsion Test Center

Trenton, New Jersey 08628

Part one : Sputter-coating and film characteristics. K. Gentner

A quarter century ago, thin, pyrolytically deposited films of re- fractory characteristics demonstrated their potential of protecting 10

atom layer thick carbon films from oxidation and abrasion. It was very

tempting, to utilize this technique for the protection of mechanical

components and structures operating in severe environments. However,

there were several factors involved which could have adversely affected

the basic structure of a metallic component. One major factor was the

relatively high temperature, and another one, the presence of reactive

atomic species.

The deposition of erosion resistant films on metals in a "clean"

processing environment at temperatures which are harmless to the

structure of the base metal can now be performed by several vacuum

deposition methods. Of these, sputtering techniques offer the largest

versatility of processing parameters to achieve extraordinary film

characteristics and selective local and sequential variations of the

material composition and film thickness at a minimum of intricate

processing fixtures or operational effort. In addition, there is

evidence, that the structural charactersitics of a mechanical component

can be substantially improved while the protective film is being

deposited.

117
Basically, a sputtering system consists of a vacuum chamber, which

houses a sputtering electrode and the mechanical component, a highvoltage power supply, and a tank of inert gas, usually argon. The

sputtering electrode, or sputtering target, can be made from one or several units of different material composition, and can have any size,

shape or spatial orientation. Once installed, a sputtering target re- quires no further attention throughout its usually long life. Process

control is simply a matter of setting a certain gas pressure in the low

micron range, and a negative potential of several kilovolts on the

sputtering target. While the vacuum chamber is usually on ground

potential, representing the anode, the work load can be held on floating or any desired potential. This optional potential of the work load can be utilized to direct a selected flux of material to and from the

mechanical component. First, by applying a high negative potential to

the component prior to the coating operation, any undesirable surface

layers - such as oxides - can be effectively removed. This way, a finely textured, sputter-etched surface structure can be prepared,

(Figure 1) which is highly desirable to achieve a good bond between

base material and film. Then, during the coating operation a proper

negative bias potential at the component can function as a very powerful process parameter to create a film composition and structure of

extraordinary properties. With the same ease and effectiveness as the

electrical control, reactive gases or vapors can be added to the sputtering gas for further variations of the film material composition. All

these variations of the sputtering operation can be easily and precisely

programmed, and once the vacuum chamber has been loaded and pumped down, no special operator attention will be required until the coating process

is completed.

We have investigated the performance of thin, sputtered coatings in an accelerated dynamic erosion test. First stage compressor blades of

the T-58 engine were first sputter-etched, and then coated inside a hollow sputtering target, fabricated at NAVAIRDEVCEN . Films of silicon

118
carbide and tungsten carbide, both blended with stainless steel, type

304, were deposited at various field configurations. The film thickness

was varied from approximately one to ten microns, with the majority of

the films at the five micron range. Sixteen differently coated blades

were prepared for the erosion test.

The test results were quite encouraging, considering the fact, that

this was an exploratory experiment to observe such factors as minimum

protective film thickness, different film composition, various deposition techniques and area distribution of erosive forces, and, furthermore, that no previous efforts were made to optimize the most favorable

coating conditions or material combinations through trial coatings or screening tests.

Visual inspection and weight loss rates of the sixteen different

coatings revealed that the erosive attack on the leading edge and con- cave surface of the blades was at least ten times more severe than that on the convex surface, which was facing the engine intake. Increasing

stainless steel content of the coatings reduced the erosion resistance

(Figure 2) , whereas an increasing negative bias potential at the blades

substantially reinforced the erosion resistance of both, the coating and

the underlying blade material. The films on the convex surface of the

blades retained for the most part their originally smooth appearance

(Figures 2 and 3), and the concave surface, having the coating abraded,

exhibited a very fine and even texture, expecially on those blades which

were sputtered at a relatively high negative bias potential (Figure 4).

In comparison to uncoated blades, or blades with much thicl'^r vapordeposited coatings (Figures 5 through 7), the sputter-coated blades ex- hibited throughout the test an extremely smooth surface, which is highly

desirable from an aerodynamic standpoint. Our next goal is, to develop new material compositions, and to advance sputtering techniques focusing

on special protection of critical areas and on obtaining overall higher

deposition rates.

119
While our main efforts are directed towards protection of new aircraft parts, the potential of sputtering techniques for repair of worn

mechanical components is being explored. An example is shown in

Figure 8„

120
121
122

124

126
VAPOR DEPOSITED BIAS SPUTTERED

Figure 7. T-58 compressor blades after 16% hours erosion test

(concave surface) .

127
C-22 C-22

1 INCH

BEFORE

1 INCH

AFTER

Figure 8. Used T-58 compressor blades, before and after sputtercoating (concave surface)

128
PART TWO: EROSION AND THE GAS TURBINE ENGINE

J. E. Newhart

Naval Air Propulsion Test Center

Trenton, New Jersey 08628

INTRODUCTION

High velocity impact of solid particles such as sand or dust, can

cause significant erosion damage to aircraft propulsion components in

helicopters, hovercraft, and V/STOL aircraft. The propulsion systems

are normally highly optimized to obtain the power to weight ratio or

thrust to weight ratio required for V/STOL aircraft and thus more suscep

tible to performance degradation from erosion. Operation of small gas

turbine powered helicopters, e.g., over dusty unimproved land areas routinely ingest up to 25 pounds of sand and dust for every hour of low altitude operation. Consequently, many billions of high- velocity , abrasive dust particles impact the critical airfoil surfaces of compressor

blades and vanes eroding away the thin metal tips and trailing edges.

The inevitable result is progressive and rapid degradation of engine per

formance, often to the extent of premature engine failure within 10-20

percent of the normal overhaul time. The accelerated replacement of ero

sion damaged helicopter turbines in Southeast Asia was estimated to cost

the U.S. Government about $150 million a year.

Two approaches can be used to combat this problem. One is to develop

a filtration system that will remove most of the solid particles from th

air stream. The other is to design blading materials and/or blading pro

tection that are not appreciably affected by solid particle erosion.

Dust filters and separators invariably decrease engine efficiency, in129
crease critical weight tolerances, and require constant maintenance.

Therefore, the optimum solution would appear to be new materials protection concepts' with significantly improved innate erosion resistance.

CONCEPTS OF EROSION

Previous investigators have postulated equations of motion for angular

dust particles striking a solid target (Ref. 1). For ductile metals and

alloys, the majority of investigators have assumed that a micromachining

process is responsible for erosion - a sharp corner of the impacting particle acting as a single-point machine tool. For hard, strongly elastic

alloys, borides, carbides, other ceramics and intermetallics , (materials

with negligible or limited plasticity), it has been assumed that erosion

occurs by random surface cracking, joining of crack networks, and subsequent spalling. These equations (Fig. 1) predict that average erosion

loss per particle should be proportional to the particles kinetic energy

2 (Mu ) and inversely proportional to the minimum flow or yield strength

for ductile targets. The usefulness of the empirically derived equation

for elastic materials is more limited because the particle velocity exponent varies from ^3. 3 to 6.5 (not a simple particle energy relation)

depending upon the material class. However, for a given elastic material,

erosion is predicted to be inversely proportional to the material's ultimate strength. For the best empirical curve fitting, the erosion losses

are also made a function of impingement angle, a, inasmuch as maximum

erosion normally occurs at some intermediate angle between 9 and 90 degrees (typically 20t30 deg.) for ductile materials; and at ~90 degrees

for elastic materials (Fig. 2).

130
PARTICLE OF MASS (M) AND VELOCITY (u)

VOLUME OF MATERIAL REMOVED (Ey)

Ev -Kf(a)^i

Mu

FOR DUCTILE MATERIALS

3.3 ,o 6.5 p0R INELASTIC MATERIALS

where o - ainimuM flow or ultimate strength of the target material

K a constant for the specific erosion system

f(o)« function of incidence angle

M * dust particle mass

" dust particle velocity

FIGURE 1 EQUATIONS OF MOTION FOR DUST EROSION

- //

FIGURE 2

Glass

V

\ \

SAE 1020 Steel N Copper

Al u mi num

Theoretical

30 INCIDENCE ANGLE ( deg

)

60 90

EFFECT OF IMPINGEMENT ANGLE ON DUCTILE AND ELASTIC

MODES OF EROSION

131
TURBINES VS. DUSTY ENVIRONMENTS

Measurements of dust concentrations in clouds formed by an H-21 helicopter in stationary hovering (1- to 10-foot heights) at Yuma, Arizona,

and Ft. Benning, Georgia, have yielded average concentrations ranging

3 3 from 2 to 4 mg/ft (average) near the rotor hub to 12 to 19 mg/ft (average) near the area of rotor blade overlap (Ref. 2). Dust concentrations

increased by a factor of 3 during takeoff and landing maneuvers, and by a

factor of 5 with another helicopter operating nearby. It should be noted

also that the ingested dust is further concentrated within the compressor

by centrifugal action, which "throws a majority of the sand to the outer

portion of the annulus, resulting in extremely high concentrations" near

the blade tips (Ref. 3). Sand from these sites is normally high in silica, with an average specific gravity of 2.4 to 2.6. Typical particle

size distribution of airborne dusts were: 98 to 100 percent <200u; 40 to

60 percent <100u; 15 to 40 percent <50u; and 5 to 10 percent<15u. Particles larger than 500u were not airborne (Ref. 2).

Of the quantitative erosion studies upon actual turbines, some of the

most relevant field work was conducted by Montgomery and Clark (Ref. 4),

who determined erosion rates as a function of dust particle size and concentration for a Solar Mars gas turbine. This study showed that within

limits of repeatability, the erosion factor

E grams weight loss

grams dust ingested

3

appears constant for any realistic dust concentration up to 13 mg/ft .

132
From this information they were able to predict the turbine life as being

proportional to a constant, which is representative of the engine type,

and inversely proportional to the product of the particle size according

to the relation

A = 1.38 L =

C x d C x d

m m

where L = engine life in hours

A = constant (1.38 for the Mars engine)

3

C = dust concentration in mg/ft

d = maximum particle size in microns

m

It was reasoned from this relation that all particles above 2 to 3 microns should be filtered from the air stream to ensure an acceptable turbine life expectancy for high dust concentration areas. Bianchini and

Koschman of Allison also carried out field studies of dust erosion in

the T63 turbine using 0 to 200-micron (Arizona Road (AR) Dust, coarse)

dust particles, initially at 15 mg/ft (Ref. 5). They found that engine

power decreased almost linearly with the total dust ingested, and that

this function was apparently independent of the dust concentration (Figure 3). Also, they noted that particles up to 10 microns in size had

only a small effect on engine power. On the other hand, the erosion

effects of 0-80 micron dust (AR fine) and 0-200 micron dust (AR coarse)

were severe, but not discernibly different.

133
100

90

80

70

A-' pi 60

o

p_, H

40

30

Field Tests: Tuma, Arizona —Controlled Tests: Allison Dirision

1 1 1 r

0-80 Micron AR Fine Dust at 0.0035 gram/ft

O 0-200 Micron AR Coarse Dust at 0.01U gram/ft • 0-200 Micron AR Coarse Dust at 0.C015 gram/ft 3

A Contaminant Ingested Hovering Over "Silt" and

"Sand Test Sites Yuma, Arizona

i

4 6 8 10

TOTAL CONTAMINANT INGESTED (lb)

12 14

FIGURE 3

T63-A-5A Engine Sand and Dust Ingestion Tests; Comparison of

Erosion Effects From Field Tests and Controlled Ingestion Tests.

Rapp and Rosenthal of General Electric (Ref. 3) also found that power degradation and the number of safe landings could be directly related to

the pounds of dust ingested. Compressor blade and vane erosion was severe on engines without inertial filters (Southeast Asian operations).

As a general rule, large particles caused substantial damage to the first

stage blades, all along the leading edges. Large particles were apparently shattered by impact with the first stage blades and accelerated to

high velocities through the latter stages, so that erosion past the first

stage proved similar regardless of initial particle size ingested. Be13b
cause of centrifugal effects, the site of blade damage was progressively

transferred to the blade tips and trailing edges, the higher the stage.

To counteract erosion, General Electric experimented with hard blade

coatings, the most successful of which were an electro-deposited hard Cr

plate and a beryllium diffusion coating. Embrittling side effects prevented their general application to AM-355 and similar stainless steel

alloys blades, however (Ref. 3).

A difference of opinion exists over the effectiveness of small dust particles (^15 microns) in eroding turbine components. Mr. J. M. Clark, Jr.,

of the Southwest Research Institute cites data which shows that, in 660

3 hours of ingesting 0-5 micron dust at a concentration of only 0.33 mg/ft ,

a gas turbine was destroyed as an operational unit. He discounted arguments that very fine dust particles were decelerated or deflected significantly by the boundary layer on airfoil surfaces. G. Thomas, also of

Southwest Research Institute, has published data showing that a Solar

turbine (Model T-221N-1) was erosion damaged sufficiently to become inoperable after running 132 hours with an inertial separator on the air

inlet. Ninety percent of the dust ingested into the turbine was determined to be ^14 microns. Based upon other data, engine manufacturers

such as General Electric and Allison feel that engine erosion is a minor

problem with particles smaller than 10-15 microns. Hence, they stress

that properly maintained inertial separators and mechanical filters

should be adequate to effectively nullify the erosion problem. (The

inertial filter has a marked cutoff in efficiency somewhere between 10-15

135
microns.)

The conclusions to be drawn from the published data regarding the effect

of ingested dust on gas turbine performance are summarized as follows:

Serious erosion of gas turbine components will be caused by

ingestion of airborne dust generated by average military operations in unimproved areas. Efficient air filters will be required on all gas turbines to remove dust particles above 2 to

3 microns in diameter if improved erosion-resistant materials

are not available.

Anticipated gas turbine life varies inversely as the product

of the maximum particle size and the dust concentration, and

the erosion factor is independent of the rate of ingestion (i.

e., dust concentration).

The total erosion is independent of dust concentration within

realistic limits, and is a direct function of the total weight

of ingested dust of any given particle size for a given gas

turbine. In the particle size range of airborne particles,

there is negligible erosion of the inducer section of the compressor impeller or other areas where the absolute velocity of

the entrained dust is low.

A simple cycle gas turbine demonstrates little or no external

evidence (by observation or readout) of complete destruction

by dust erosion prior to the point of complete failure. Vibration of gas turbines does not appear to increase until

136
almost total destruction of the rotating components by dust erosion has

occurred

.

Field studies give only a generalized or end view of the erosion phenomena, with limited potential for insight into the basic mechanisms of ero

sion. This is because erosion in a turbine occurs over a wide and chang

ing range of conditions that are difficult to control directly as variables .

EROSION TESTING

Recognizing the fact that erosion in a turbine is a product of many

variables, then the development of a meaningful "bench" type test becomes monumental. Table 1 lists most of the variables that must be considered, incorporated, and controlled. It appears as if a full scale

turbine engine is the only answer as a test vehicle; however, the cost

consideration precludes this use as a routine test. Two of the more

widely used erosion testing devices are the Roberts Jetabrader and the

S.S. White erosion test. These are small, self-contained, primarily research oriented testing equipment. The Solar Division of International

Harvester has developed a new erosion test facility which can apply and

evaluate most of the variables in Table 1 (Ref. 6). The primary drawback of all these test techniques is the inability to evaluate more than

one specimen (material and/or coating) per test.

137
EROSION TEST VARIABLES

Test Dusts

Particle Size

Particle Velocity

Impingement Angle

Dust Concentration

Carrier Gas

Carrier Gas Pressure

Carrier Gas Temp

Target Materials

(Table 1)

Arizona Road Dust (natural desert sand 68% SiO^)

Laterite (35% Si0

2

)

Pure Silica Sand (100% Si0

2

)

Pure Alumina (100% A1

2 0

3

)

Up to 200 microns diameter

100 to 1500 fps

20 to 90 degrees

3

5 to 200 mg/ft"

Air, pure or mixed gases

1 to 16 atmospheres

Ambient to 800°F

AM 355, 410 stainless, Ti-6A1-4V, 17-4PH, 17-7PH

various coatings

The Naval Air Propulsion Test Center (NAPTC) has designed and operates

a dynamic erosion facility. The test rig consists of a standard T58-GE10 engine first stage rotor, front frame, bullet nose, and bellmouth

(Fig. 4). The rotor contains 30 blades of cast AM-355, the standard

alloy used in current model engines. An inlet cloud chamber, close-coupled to the bellmouth (Fig. 5), allows a uniform, dispersed flow of measured abradant to enter the bellmouth. A calibrated feeder supplied the

abradant through a funnel ejector to the inlet cloud chamber (Fig. 6).

The abradant used for all testing to date is prepared Arizona Coarse

Grade Road Dust, 0 to 200 microns, from the AC Spark Plug Division, GMC.

138
Dust concentration was delivered at 0.01 gm/ft . The air flow was maintained at 13 lb/sec (T58 requirement) and the average rotor speed was

measured at 11,000 rpm. The length of a test has been established at

15 hours for AM-355 blades, which results in a 1 gram erosion loss of

material from unprotected blades. In addition to weight loss determinations, measurements are recorded for losses in chord width (tip, midspan,

root) .

Although this coordinated program of "sputtered" coating research and

dynamic erosion test development is still in its infancy, the composite

results are gratifying. As described by the author of Part 1 of this

paper, distinct trends are shown for various concepts in the sputtered

coating depositions. These trends are integrally related to the reproducibility of the test. For example, the weight loss rate, mg/hr, for

standard uncoated blades during test shows an average deviation of less

than 3 mg/hr.

A summary of the erosion protection development program at this stage

shows the following:

1. Sputtered coatings -

a. Can be applied relatively "cold" with minimum effect to subs trates

.

b. Improve aerodynamic properties of airfoils and do not degrade

these properties with erosion loss.

c. With controlled bias potentials, reinforcement to erosion

resistance is obtained for both the coating and substrate.

139
2. Dynamic erosion testing

a. Evaluates actual hardware indicative of service experience.

b. Abradant injection technique produces random impingement angles of random particle sizes.

c. Produces a true picture of erosion prior to channeling and

fracture of the abradant.

References

1. Finnie, I., "Erosion of Surfaces by Solid Particles", Wear, Vol. 3,

pp 87 - 103

2. Rogers, S. J., Evaluation of the Dust Cloud Generated by Helicopter

Rotor Downwash", USAAVLABS T.R. No. 67 - 81 of March 1968

3. Rapp, G. C. , and Rosenthal, S. H. , "Problems and Solutions for Sand

Environment Operation of Helicopter Gas Turbines", ASME Publication No.

68-GT-37

4. Montgomery, J. E., and Clark, J. M. , Jr., "Dust Erosion Parameters

for a Gas Turbine", SAE Meeting, Atlantic City, N.J. June 1962

5. Bianchini, V., and Koschman, R. B. , "T63 Engine Sand and Dust Tolerance", Report No. 66-ENV-18 of Allison Division, General Motors Corp.

6. Smeltzer, C. E., Gulden, M. E., McElmory, S. S., Compton, W. A.,

"Mechanisms of Sand and Dust Erosion in Gas Turbine Engines", USAAVLABS

Tech. Report 70-36 of August 1970

iko
141
;
143
DISCUSSION

G. F. Schmitt, Jr., Wright-Patterson Air Force Base : Was there some particular reason for settling for a coating thickness of 10 microns

or less? Admittedly you have pretty close tolerances on the thicknesses of coatings that can be put on these blades because of aero- dynamics, etc. The coating thickness is certainly not limited to

anything as small as that. Is there not some reason to believe that

even better protection might be obtained with a somewhat thicker

coating?

K . Gentner : The reason I used only up to 10 microns was that I wanted

to see definitely any erosion damage on the films so I could learn

something. It showed that on the back sides of the blades, the film

was eroded away before we came to the end. By using a sputtering

process I can coat selective coatings over the area. That means I can give the front side, which needs very little protection, a very

light coating, and give the back side a real heavy coating. It has

been shown that the back side erodes at least 10 times more severely

than the front. Engine people are interested in a minimum added

weight from a coating. With sputtering, a light coating can be put

on the front side and a heavy coating on the back side which keeps

the added weight to a minimum.

J. E. Newhart : The thicker the coating, the greater are the fatigue

losses in the substrate. At about a 0.0015 in. thickness and with

certain types of high temperature coatings, a 40% loss in the fatigue

strength of the substrate can be expected. In some turbine engines

there isn't much of a pad to play with.

G. F. Schmitt, Jr.You just answered my next question which was about the effect on the fatigue life. I recognize that sputtering is a pretty low temperature process as opposed to vapor deposition.

K. Gentner: Actually, to have a high sputtering rate, the sputtering

electrode is chilled. The hotter the electrode, the lower is the

sputtering rate. If a film is deposited, it should be annealed on

the plate. In any other process a high temperature is used which is

not good for the base plate, but in sputtering the energy comes from

the high speed of the incoming particles. That means the blade can

be kept relatively cold and all the energy to anneal the film comes from the incoming particles.

G. F. Schmitt, Jr. : What about using it on titanium? You said you

sputtered titanium carbide on titanium. What about anything else and

what sort of adhesion or lack thereof would you expect?

Ikk
K. Gentner : High adhesion is obtained if all the oxides are removed

first. How much material should be removed? I have in many cases not removed more than 1 micron. I went as high as 4 microns in one case but it was not necessary.

R. L. Johnson, NASA, Lewis Research Center: We've used the processes

of sputtering and ion plating primarily for lubricating coatings for

sometime. A year and a half ago there was a conference at Lewis

that dealt with both sputtering and ion plating and if anyone is

interested in it I can provide copies of this . I think there is a point here that needs to be raised, though. The question was brought

out in the discussion of the effects on fatigue. The usual concepts

of the adverse influence of coatings on fatigue don't apply with

these types of coatings. We have been doing bending fatigue studies

with metal specimens with coated surfaces, and we find in some of

the surface analyses that sputtered or ion plated materials put on with reasonable processes will, in fact, immobilize surface defects,

rather than create them which occurs with some other kinds of

coatings. We've shown an increase of 10-50% in endurance life in

bending fatigue studies on thin flat sheets. This is with only

about 2000 angstrom units thickness of metal coated on another metal.

I think this matter of fatigue effects can be advantageous. I mentioned earlier that these coatings will, in effect, tie down

surface defects. Ion plated metals will penetrate grain boundaries,

and although we haven't experimented with them, in fluid erosion

studies, I think that they can mitigate against some of the mechanisms

that have been discussed here earlier where slip is occuring at grain

boundaries, etc. We had some erosion problems with a fluid in a hydraulic system pump with a bronze surface. The problem was associated

with the fluid itself, but we tried to salvage some parts using sputtering of monel. There was erosion damage on the operating parts of

a pump similar to the hydraulic pump, the piston pump that we saw here earlier, in a matter of a 1/2 hour with the bronze. With the

monel coating, which was a few thousandths of an inch thick, much

thicker than we usually use, it took at least 10 hours before we saw

any sign of damage. So the coatings can be used in this area without

going to the extremely hard coatings that are being talked about. The

process is a very flexible one.

lU-5
RAIN DROPLET EROSION MECHANISMS IN TRANSPARENT

PLASTIC MATERIALS

George F. Schmitt, Jr.

Air Force Materials LaboratoryElastomers & Coatings Branch

Wright -Patter son Air Force Base, Ohio ^5^33

The damaging effects of rain erosion on optically transparent

materials such as are used for windshields and infrared, television and

laser guided missiles have become an increasingly severe problem as

the number of systems using this type of guidance have increased. The

rain erosion of these materials is a twofold problem; first of all, subsonic erosion and runway debris damage during captive carry beneath

the wings on the aircraft, and secondly supersonic erosion during free

flight after being launched.

On glass materials the subsonic erosion damage takes the form of

isolated pits which eventually join together with subsequent chunking of

larger pieces. At supersonic speeds catastrophic fracture occurs with

large scale cracking and material loss. All materials suffer a significant deterioration (up to 80% loss in transmission) after rain erosion

exposure even when no measurable weight loss has occurred.

The front end covers for optically guided missiles called "domes"

are typically constructed in hemispherical shapes to enable optimum

transmission through them. However, this shape with its large

normal (90 ) or near normal incidence angles is the worst geometry

possible from an erosion standpoint. Reduction in impingement angle

is always desirable to reduce erosion effects but this is not compatible

with the tracking requirements.

A potential important use of transparent plastic materials which

possess some degree of subsonic rain erosion resistance and a moderate degree of thermal capability for withstanding aerodynamic heating

during captive carry on supersonic aircraft and after launch is for guided missile optical covers. Polycarbonate and polysulfone plastics have

improved thermal stability over acrylics, lower cost than glass, excellent strength properties, and good visual transmission characteristics.

The purpose of this investigation v as to determine the effects of subsonic

rain erosion exposure on the transmission of radiation at a single wavelength ( 1 . 06H-) through these plastics.

li+6
EXPERIMENTAL APPARATUS

The erosion investigations of these transparent plastic materials

were conducted by the Air Force Materials Laboratory on a rotating

arm apparatus (see Figure 1) . This equipment includes an 8 -foot

diameter propeller blade made of 4340 steel mounted horizontally and

powered by a 400-horsepower electric motor. It is capable of attaining

variable speeds up to 900 mph at the blade tip where the specimens are

inserted. A detailed description may be found in Reference 1.

The speed of the equipment is regulated by a thyristor power supply

from which rigid control is possible. A revolution counter is utilized

for monitoring velocity; vibration pickups are used for gauging specimen

balance and smoothness of operation. The rotating specimens are

observed using a closed circuit television camera and a stroboscopic

unit synchronized with the blade revolution.

Mounted above the blade is the water system used to simulate the

rain environment. The 8 -foot diameter, 1 inch aluminum pipe ring is

equipped with 96 equally spaced hypodermic needles to yield a rainfall

simulation of 1 inch per hour. The hypodermic needles are No. 27

gauge which produces rain droplets of 1 . 5 to Z . 0 mm diameter as determined photographically. The water system operated with low

pressure in the spray ring enables a stream of undistorted water drops

to impinge on the material specimens.

The rotating arm simulation apparatus is used extensively in erosion investigations because of the correlation between it and actual flight

test results. Various materials were evaluated under actual flight conditions in rain utilizing an F-100F aircraft. These tests verified that

the rankings and modes of failure obtained on the rotating arm are borne

out in actual flight experience.

Flat panel specimens were run with rain impingement at discrete

angles of 90 , 60 , 45° and 30° with respect to the specimen surface.

Materials

The materials for these experiments were Lexan 9034, UV stabilized polycarbonate from General Electric Company and Bakelite P-1700

polysulfone from Union Carbide Corporation. Typical vendor -supplied

physical properties are shown in Table I. Polycarbonate and polysulfone are candidate materials because

their spectra shows transmission of 90.4% and 88.2% respectively at

10, 000 cm"^; i. e. , the wavelength, 1.06 microns, of interest.

Impingement and Transmission Experiments

For laser transmission, the DK-2 recording spectrophotometer was

used with a beam size 1/4 by 1/4 inch at 1.06 microns . With the

sample removed, the chart pen is adjusted to 100 percent. The shutters

are closed and the chart pen is set on zero. The shutters are again

147
opened, the sample is positioned in the appropriate spot normal to the

beam direction, and the total luminous transmittance is given by the

chart pen.

Nominal velocities of 400, 500 and 600 MPH and rainfall conditions

of 1 inch/hour simulated rainfall (1. 8 mm dia drops) were used.

Weight loss measurements were attempted but in all but nine of 110

samples a weight gain occurred. The weight gain represented approximately 0. 1% to 0. 2% of the total sample weight. It was approximately

the same for both materials and did not vary significantly as a function

of velocity, impingement angle, or exposure time. These results indicate that significant material does not have to be removed from the surface to obtain a decrease in optical transmission. As suggested by

Figures 2, 3 and 4, the polycarbonate or polysulfone surface layers

tend to pit under the impact of the rain drops. This pitting causes

hazing or frosting of the surface.

The difference in the detailed character of the eroded polycarbonate

surface compared to the polysulfone is shown in Figures 2 and 5. In

Figure 2, a polycarbonate specimen which was exposed at 400 MPH for

30 minutes exhibits individual pits which are somewhat isolated from

each other. The polysulfone in Figure 3 after similar exposure shows

uniform light surface frosting (pits on a much smaller scale) . It is this

difference in the form of damage which explains the polycarbonate's

better retention of transmittance after rain exposure (i. e. , through the

relatively undamaged areas) whereas the polysulfone' s surface would

diffract a larger percentage of the energy and result in lower transmission.

The progression of damage in polycarbonate is shown in Figure 4

for specimens after 90 minutes at 400 MPH. With this long term

exposure, the pits have grown and intersected to obscure the surface.

The tests were terminated during the incubation period of the

erosion process. This is the time during which significant cumulative

damage is occurring on the materials but no measurable material loss

has yet occurred. The decrease in transmission with no weight loss is a highly significant point because the short exposures on the rotating

arm in these tests would correspond to very limited exposure on the

missile dome during captive carry and yet up to 50 percent decrease in

transmission has been measured.

An incubation period associated with weight loss for transparent

materials does not appear to be an appropriate measure of the onset of

significant damage. A more appropriate criterion for ranking of these

materials would be the point (exposure time or number of impacts) at

which a particular percentage of the original transmittance has been

lost. This coituld be varied depending upon the sensitivity of the optical

system to this degradation.

1U8
Total transmission measurements at 1. 06 microns wavelength are

given in Table II. Where two samples were run at each condition, the

results are averaged. The data in Table II is summarized in Figure 5 for polycarbonate and Figure 6 for polysulfone where transmission as a percentage of the initial (prior to exposure) is plotted versus time of

exposure in one inch per hour simulated rainfall. The parameters in

this plot are the normal components of the velocity (v sin 0) corresponding to 300, 400, 500 and 600 MPH. As can be seen, only at a normal

component of velocity of 500 MPH does significant decrease in transmission through the polycarbonate occur. By contrast, even a combination as low as normal velocity equal to 400 MPH exposure results in up

to 20% decrease in transmittance through the polysulfone material.

DISCUSSION

The importance of the effects of erosion on reduction in transmission during the incubation period cannot be overemphasized. Traditionally, the portion of the erosion process which has been of most concern has been the accumulation period when mass loss occurs as a function of continued exposure to the environment. The loss of material

with its consequent reduction in performance of the system was the

major problem (see several papers each in References 2 and 3) . However, in the case of optical materials the effects occur before

any measurable mass loss begins. Furthermore the appearance of the

surface of the transparent materials after this limited exposure would

normally be considered slight erosion or little effect. Yet it has

seriously degraded the most important property of the materials; namely

their ability to transmit the radiation. The obvious effect this has on

knowing when a degraded dome is no longer operative and must be replaced from a maintenance standpoint is potentially serious.

The influence of physical properties on the erosion behavior is not

easily discernible. As can be seen in Table I the physical strength

properties of the polycarbonate and polysulfone plastics are so similar

that little difference in erosion performance might be anticipated. The

major difference in strength between these two materials is their impact

properties as measured in a notched Izod test. The erosion resistance

of most materials cannot always be directly correlated with impact

strength but is usually quite good (Reference 4) . The order of magnitude difference between the polycarbonate (16 ft-lb/in) and the polysulfone (1.3 ft-lb/in) would provide a difference in response to the droplet

impact.

An extensive discussion of the correlation between rain erosion

resistance of partially crystalline plastic materials and their notch

impact strength has been presented by Oberst (Reference 4) . He shows

a general correlation of increasing erosion resistance with increasing

impact strength for polymethylmethacrylate, polycarbonate, polyacetal

149
and polyethylene materials. Polysulfone is not included, however.

Furthermore, this correlation is not supported by specific experimentation to test it but is based on single experimental erosion values of Hoff

and Langbein of Dornier Systems.

Adler, Morris and Wahl (Reference 5) specifically investigated a number of bulk plastics of varying degrees of crystallinity , including

polyur ethane, polyethylene, polycarbonate, acrylic polyvinyl -chloride,

polyphenylene oxide, acetal and nylon and determined that no correlation could be drawn between notch impact strength and rain erosion

resistance.

As shown in Table I, the polycarbonate and polysulfone are so close

in hardness (even though this is a macromeasur ement) that no correlation can be drawn.

Previous investigation of thermosetting and thermoplastic resins

has shown that the ductility of the plastic has a significant influence on

its erosion behavior (Reference 6) . The erosion resistance and resulting optical transmittance appears to be a function of the ductility of the

plastic also. As indicated in Table I, the elongation of the polycarbonate

is 110% compared to 50% for the polysulfone. The morphology of the

polycarbonate surface after erosion for 30 minutes at 400 MPH (see

Figure 3) , exhibits no undamaged area with incipient fine scale pitting

(or frosting) over the entire surface. This behavior is certainly in

keeping with the better retention of transmission by the polycarbonate

compared to the polysulfone.

Ductility is a function of temperature in the plastic of course, and

the higher heat deflection temperature of the polysulfone (345 F) compared to polycarbonate (270°F) may be detrimental in that it does not

become more ductile during the impact process when heat is generated

by energy deposition.

Although not specifically investigated in this program, polymethylmethacrylate (Plexiglas) has exhibited similar frosting type obscuration

of the surface very rapidly under identical erosive conditions. With its

very low elongation (<10%) , it exhibited extremely brittle chunking

behavior at 400, 5 00 or 600 MPH. This would further confirm the

importance of ductility for transparent plastics retention of transmittance. Obviously only a moderate amount of ductility is really desirable since too soft a material would be sufficiently deformed by the

impacting drop to distort the transmittance.

A recent theoretical development by Springer and Baxi (Reference

7) shows promise for correlating the fatigue properties of homogeneous metals and plastics to their incubation period and subsequent mass

loss behavior in subsonic rain erosive environment. Properties of the

materials required are the ultimate tensile strength, endurance limit,

Poisson's ratio, and the fatigue life. This correlation formula has

150
shown excellent agreement with all existing experimental erosion data

(mostly metals with some bulk polymers) . An adaptation of the above theory to account for the change in surface morphology during the incubation period has not been accomplished.

The theory also does not treat the case where each individual drop

impact does damage although this was considered in the analysis. Concepts which utilize this approach may provide insight into the governing

materials properties for the erosion of transparent materials.

CONCLUSIONS

1. Significant decreases of up to 80% reduction in laser transmit - tance at 1 . 06 microns wavelength through polycarbonate and polysulfone

transparent plastics are measured after short term subsonic rain

erosion exposure.

2. The major changes in transmittance of these plastics occur

during the incubation period of erosion; that is, before mass loss begins.

3. Although both are lacking in long term subsonic rain erosion

resistance, polycarbonate is superior to polysulfone in retention of

transmittance after exposure.

4. Greater notch impact strength and ductility of the bulk plastic

improves its erosion resistance and reduces subsequent loss in transmittance properties.

5. The change in transmittance through the polycarbonate and polysulfone materials can be approximately correlated with the morphology

of the eroded surface after exposure.

6. Transparent plastic materials for optical, infrared or laser

applications, are lacking in rain erosion resistance.

7. Polycarbonate and polysulfone transparent plastics have limited

application for missile dome covers where rain erosion exposure at

velocities of 500 MPH or greater are involved. However, these plastics

can be used with the proper design such as low impingement angle and a favorable environment such as short time exposure, low subsonic velocities, or low density rainfields.

REFERENCES

1. C. J. Hurley and G. F. Schmitt, Development and Calibration of a Mach 1.2 Rain Erosion Test Apparatus , AFML-TR-70-240, October

1970.

2. A. A. Fyall and R. B. King, eds. Proceedings of the Third

International Conference on Rain Erosion and Allied Phenomena , Royal

Aircraft Establishment, Farnbo rough, England, Volumes I and II, 11-

13 August 1970.

3 . Characterization and Determination of Erosion Resistance , American Society for Testing and Materials Technical Publication 474,

Philadelphia, Pa., 1970, 434 pp.

151
4. H. Oberst, Rain Erosion and Molecular Properties of Synthetic

Materials , Royal Aircraft Establishment Library Translation No. 1335,

December 1968.

5. W. F. Adler, J. W. Morris, and N. E. Wahl, private communication.

6. G. F. Schmitt, Materials Parameters that Govern the Rain

Erosion Behavior of Polymeric Coatings and Composites at Subsonic

Velocities , Presented at Third ASTM Symposium on Composites:

Testing and Design, Williamsburg, Va., 21-22 March 1973.

7. G. S. Springer and C. B. Baxi, A Model for Rain Erosion of Homo - geneous Materials , Presented at ASTM Symposium on Erosion, Wear

and Interfaces with Corrosion, Philadelphia, Pa. 26-27 June 1973.

ACKNOWLEDGEMENT

The author wishes to acknowledge the assistance of Dr. Richard C.

Hall of the Martin Marietta Corporation, Orlando, Florida who provided

the materials specimens and transmission tests.

jr>js

C

tsMo 3

152
TABLE I PHYSICAL PROPERTIES OF TRANSPARENT PLASTICS

Lexan 9034 Bakelite P1700

Polycarbonate Polysulfone

Specific Gravity 1.2 0 1.24

Tensile Strength, psi 9000 10,200

Tensile Modulus, psi 345,000 360,000

Elongation, % 110 50 - 100

Flexural Strength, psi 13,500 15,400

Flexural Modulus, psi 340,000 390,000

Shear Strength, psi 10,000

Compressive Strength, psi 12,500

Compressive Modulus, psi 345,000

Rockwell Hardness M70(R118) M69 (R120)

Izod Impact Strength 16 1.3

at 72° F, ft lb/inch-notch

Heat Deflection Temperature, °F 270 345

at 264 psi (0. 010" deflection)

All values obtained by standard ASTM tests.

153
TABLE II

TOTAL TRANSMISSION CHANGES AS A FUNCTION OF

EXPOSURE IN RAIN

Normal Velocity

Component (MPH)

Impact

Angle

(degree)

Nominal

Velocity

(MPH)

ExDOKure

Time

(min)

Percent Transmittance*

as a Function of Original

Polvcarbonate

i+oo 90 1+00 5 98.1+ 9^.7

10 98.2 89.U

20 99-1 87. U

1+33 60 500 2 98.9 9^.7

5 97.8 89.3

10 97.1 75.2

20 93.0 53.1

k2h ^5 600 2 96.9 95.1

5 97.3 88.7

10 93. U 82.6

20 88.1 72.7

500 90 500 1 99.8 9^.1

2 98.1+ 86.2

5 96.6 77.7

10 90.7 53.9

20 7I+.0 18.3

519 60 600 1 95.7 81+.

2 9^.8 8U.

5 90.0 6k. h

10 78.7 f w • 1 1+6.6 20 65.8 25.1+

600 QO 600 1 89.2 81.7

2 82 ^ ^ u • j

5 56.6 52.6

*>0 600 97.3 95.3

10 97.0 95.5

20 95.8 91.5

35^ h5 500 5 96.5 93.7

10 97.1 9h.G

20 97.2 8l+.

3U6 60 1+00 20 97-^ 91.7

250 30 500 10 97.3 95.6

20 97.9 95.5

* Measurements at 1.06p. after exposure to 1 inch/hour simulated rainfall.

Average of two specimens except for 90 specimens which are only one measurement.

15l+

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CD Pd

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156
Figure 4. Severe Pitting on Surface of Polycarbonate

after 90 minutes at 400 MPH

157
158
g o o o o o

g CD ID * C\J "1VIJLINI dO !N33d3d V SV NOISSIIAISNVH1

159
ANALYSIS OF NAVY RADOME FAILURE PROBLEMS

G. Jo Tatnall and K 0 Foulke

Naval Air Development Center

Aeronautical Electronics Technology Department

Warminster, Pennsylvania 1897^-

The analysis or diagnosis of Navy radome failure problems was done

at the request of the Commander, Naval Air Force, U.S. Pacific Fleet.

It was performed under the direction of the Materials Readiness Officer,

Carrier Division Five, Attack Carrier Striking Force, Seventh Fleet.

The survey of environmental damage and repairs included observation of

all fighter and attack aircraft then operating within the Tonkin Gulf

from the Subic Bay Naval Base. In addition it included the observation

of all aircraft assigned to the Naval Air Station, Cubi Point and those

aircraft undergoing intermediate level maintenance at the Fleet Air

Wing Repair Activity, Cubi Point.

The aircraft observed aboard three attack carriers and ashore ex- ceeded two hundred. The aircraft models observed included all first

line Navy fighter and attack aircraft. Aircraft time since overhaul

(PAR) varied widely although the majority of aircraft were well within

the second half of their PAR cycle.

These aircraft had in general been operating within an extremely

hostile environment. Combat missions into Viet Nam were continuing daily in spite of the monsoon season. A large percentage of the

flight hours each day were through moderate to heavy rainfall. In

addition, most of the preventative maintenance procedures for radomes

were being omitted because of bad weather and tight aircraft scheduling.

The task was to carefully observe all these aircraft model by model

and radome by radome and to organize the damage findings into some reasonable order. From this survey the critical problem areas were defined and a plan was developed for failure prevention.

160
As more and more aircraft were observed it became increasingly

obvious that the erosion damage was a serious problem, that many electronic systems had been degraded through poor and improper repairs, and

that certain airframe manufacturers had done a better job of providing

realistic environmentally resistant radomes than had other manufacturers.

It was further obvious that the military guideline specification, MILR-7705, "Radomes; General Specification for" was ineffective in providing adequate hardware for high performance all-weather aircraft. There

was a serious lack of materials and skills to effect satisfactory re- pairs or to provide for a reasonable degree of preventative maintenance

at the squadron level. The technicians were lacking information and

help which would assist them in doing their job more satisfactorily.

At the conclusion of the survey or diagnosis of the problems the

surveyor could not help feeling dedicated to the solution of as many of

these problems as possible. An internal Navy report, "Task Force Seventy

Seven, Aircraft Antenna Cover Rain Erosion Damage Survey" was prepared

aboard the USS ENTERPRISE, approved by the Materials Readiness staff,

signed by the Commander, Carrier Division Five and forwarded to the Commander, Naval Air Force, U.S. Pacific Fleet and to the Naval Air Systems

Command Headquarters. This report defined the causes of the problems,

listed many specific recommendations for individual product improvements

and made several general recommendations. The recommendations included

both short term improvements and long term improvements.

From the CTF-77 report came a series of quick reaction type projects

and a subsequent series of longer term projects. All projects were in

direct response to the needs of the fleet in accordance with the findings of the survey.

A major quick reaction effort was directed toward the development of an interim "Radome/Antenna Cover O-Level Repair Kit" . Within 4 months

some 160 of these kits were distributed among the Navy's attack carriers

world-wide for immediate radome repairs.

The repair kits were developed for use by personnel having no prior

161
experience with plastic repairs. The resin systems and curing processes

required no facilities or special equipment. The reinforcement materials included a low dielectric cloth which permits the increase of wall

thickness and therefore improved impact resistance on certain antenna

covers. Polyurethane tape was provided for quick and easy replacement

of worn coatings to prevent further damage to the substrate materials.

The Repair Kit Instruction Sheet outlined specific repair instructions

for the seven most chronic problems found in the survey. All of these

repairs were modeled and validated for conformance with the system and

antenna specifications.

Immediate assistance was provided for one small antenna cover by inhouse manufacture and shipment of 125 parts within 30 days. These re- placement parts were installed by squadron personnel whereas the entire

antenna assembly was previously being replaced when replacements were

available.

Materials, tools, equipment and repair drawings were shipped to the

Fleet Air Wing Repair Activity at Cubi Point to elevate their local

intermediate level radome repair capability. Development of this capability made the depot less dependent upon spares replacement, transferred some of the repair work to the shore station and permitted more

complex repairs to be made within the Seventh Fleet area. Six of the seven Naval Air Repair Facilities (NARFS) were surveyed

to inspect their radome repair capabilities. Aircraft of all types were

inspected as received, in-process and ready-for-delivery . The condition

of radomes on completed aircraft varied from excellent to poor.

Whereas it had been previously assumed at the time of the fleet sur- vey that all PAR aircraft were returned to the squadrons with radomes

completely restored, many were being returned with their radomes in a partially damaged condition. This has since been rectified to decrease

the amount of repair work at the squadron level and to effect more professional repairs.

162
Having completed most of the short term assist projects our efforts

were turned toward the longer range objectives. At this time the

NARF's overall had improved their radome re-work capability and comment

sheets were being returned by the users of interim Radome/Antenna Cover

Repair Kits. The frequency of receiving UR's (Unsatisfactory Reports) was diminishing very rapidly.

The repair kit users comment sheets were used to improve the Radome/

Antenna Cover O-Level Repair Kit before release to the Aviation Supply

for their assignment of Federal Stock Numbers, MIL-Spec packaging and

procurement. Both the basic kit and the polyester supplement are currently available from Navy stock.

A project was created for the revision of MIL-R- 7 7 05 . This specification, originally prepared in 1944, had minor revisions in 1950 and

1955. It did not provide the proper guidelines for the preparation of

detail specifications for aircraft or electronic systems in the 1970 's. Many of the problems found in the survey could be directly related to

design deficiencies which could in turn be related to the deficient MILSpec. MIL-R-7705 was therefore completely re-written with the objective

of preventing similar problems on new aircraft. In addition to updating

the electrical parameters the changes impose environmental requirements

for rain impact resistance, rain erosion resistance, hail impact and

high temperatures. It also establishes requirements for reliability,

maintainability and certain unacceptable constructions. The new specification has been approved by the Navy and is currently awaiting acceptance or comment from the Air Force and Array.

A separate project was created under CAMSI (Carrier Aircraft Maintenance Suitability Improvement) funds to develop a service repair manual. This manual, NAVAIR-01-1A-XXX is an Organizational and Depot Maintenance Instructional Manual, titled "Aircraft Radomes and Antenna

Covers " . It is intended to provide the needed guidelines for quality

repairs and preventative maintenance at all maintenance levels. It may

also be used as a training document or as a simplified technician level

163
reference on radomes. It contains general design and performance information; detailed materials, construction and repair information; repair

techniques; and electrical test methods of repaired parts. It collects

all the needed repair information into one reference, directly relates

to the Radome/Antenna Cover O-Level Repair Kit and assures a satisfactory repair when performed in accordance with the instructions.

The manual has been designed for frequent revisions when systems are added to or deleted from an aircraft or when a new aircraft is added to

the fleet inventory. It currently includes all radomes for all fighter

and attack aircraft. The manual has been released to NATSIF for editing, illustration and publication. Preliminary copies are being re- viewed by all NARFS and by select AIRPAC and AIRLANT organizations before publication.

Other long term improvements involved the complete re-design of

several small radomes which had no chance of survival without major

changes in their material or configuration. Some effort was also devoted to the development of a low cost throw-away or disposable radome concept. Through the use of low loss impact resistant moldable thermoplastic materials it appears cost effective to replace rather than to

repair certain small radomes. Two new Navy aircraft antenna installations have already adopted this approach.

The rain erosion damage survey established a strong justification

for the need of qualification testing of materials and designs which

may have questionable life expectancy on the aircraft. Such MTBF and

damage rate testing should be required in future specifications. This

can be accomplished only if suitable test facilities and test methods

are available. Such a capability is currently being developed.

To date the total effort has proven most rewarding when measured by

the number of urgent messages, UR's and phone calls received. A true

reading for the improved level of failures can be made only by comparison

before and after this work using the 3M (Maintenance, Materials, and

Management) computer system to search out improvements in replacement

rates and maintenance man hours.

161+
DISCUSSION

G. F. Schmitt, Jr., Wright -Patterson Air Force Base : Does that

radome damage repair kit have a Federal Stock Number?

K. Foulke : Yes. It is available from the Federal Stock. It can

be ordered now.

3SiIt 3Sr

165
GOVERNMENT- INDUS TRY DATA EXCHANGE PROGRAM

(GIDEP)

EDWIN T. RICHARDS

GIDEP ADMINISTRATION OFFICE

NAVAL FLEET MISSILE SYSTEMS ANALYSIS AND EVALUATION GROUP

CORONA, CALIFORNIA 91720

The Government -Indus try Data Exchange Program (GIDEP) was originated

Ln 1959 by the Army, Navy, and Air Force Ballistic Missile Agencies.

Cnown at that time as IDEP—Interservice Data Exchange Program—its in-

:ent was to eliminate duplicate testing of parts and components by disseminating pertinent test data among Department of Defense contractors

md various government agencies.

In 1966 both the National Aeronautics and Space Administration and

:he Canadian Military Electronics Standards Agency (CAMESA) recognized

:he value of the data provided by the program and became participants,

today, GIDEP provides the interchange of specialized technical data to

ill the military services, participating government contractors and numerous government agencies such as the Atomic Energy Commission, Federal

Aviation Administration, Defense Supply Agency, and the Small Business

Administration.

GIDEP operates under a charter agreed upon by the Army and Navy MaterLei Commands, the Air Force Systems and Logistics Commands and NASA. The

J rogram Manager, organizationally located within the Department of the

tfavy, is responsible for policies and procedures, both national and inter

lational. In addition, he is responsible for establishing funding re- quirements and reporting on the cost effectiveness of the program. The

3IDEP Administration Office is the operational arm of the program. Working directly under, and responsible to, the Program Manager, the Adminis-

:ration Office maintains the GIDEP data banks and is responsible for all

jperational phases of the program. The GIDEP program management team

Includes government and industry advisory groups, which are made up of

representatives from government agencies and industry participants. They

act in an advisory capacity, presenting the requirements of the particiDants to the Program Manager and the Administration Office.

The objectives of the GIDEP program are to: reduce or eliminate dupli

:ate expenditures for development parts and components; increase the confidence level in the reliability of systems utilizing these parts and

:omponents; expedite research and development projects by avoiding repetition of tests previously accomplished; assist in the preparation of nore realistic proposals; standardize procedures for reporting test

166
information; encourage direct inter-contractor communications among technical personnel working on related projects; generate information for an alternate source of parts through more dependable data; create a general

source for test data; and provide for the exchange of test equipment

calibration procedures and related data. Mechanical and electronic failure prevention and warning of potential problems are part of GIDEP's

operational system.

GIDEP data are contained in four separate information banks: the

Engineering Data Bank, the Failure Experience Data Bank, the Failure

Rate Data Bank, and the Metrology Data Bank. No classified or company

proprietary data is included.

The Engineering Data Bank contains primarily laboratory data relating

to parts, components and materials. These data cover Qualification and

Environmental Testing, Research and Development, Evaluation Reports, and

other meaningful engineering data such as non-standard parts justification, test planning, and manufacturing processes.

The Failure Experience Data Bank contains failure data and failure

analysis reports from the field, laboratory and production line. It also

contains all of the information from the Air Force Defective Parts and

Components Control Program which contains failure analysis data by com- ponent types derived from an operation and maintenance level.

Part of the Failure Experience Data Bank is an important program known

as the ALERT system. The ALERT is a highly effective means of rapidly

providing all participants with information concerning an actual or potential problem involving a part, material, test equipment, process or safety problem. Any participant who finds a situation which he feels

to be of general concern to other participants fills out an ALERT form.

The ALERT form is submitted to the Administration Office where it is

reviewed and distributed to all participants as an ALERT. Generally,

this process is completed within twenty-four hours.

The ALERT system may be used for any type of pertinent information

relating to any of the data banks. It is issued to identify such items

as faulty design, faulty test equipment or calibration procedures, or other production and processing problems. The advantage of a short notification time in the ALERT system provides cost savings to participants

by avoiding the repetition of problems which have been previously identified or solved.

The Failure Rate Data Bank contains field performance data relating to

parts, components and materials. Detailed information concerning failure

rates, stress levels, mean time to repair, level of test specification,

failure mode, test environment and other pertinent information is contained in the Failure Rate Data Bank.

I67
The Metrology Data Bank contains calibration procedures and general

information on test equipment. Calibration procedures prepared by both

the military and industry, covering most electrical and mechanical test

equipment, are available to participants. In addition, under a program

called SETE (Secretariat for Electronic Test Equipment) , other types of

information such as test equipment evaluation reports are available.

An additional function of GIDEP is an International Data Exchange

Program called EXACT. Functioning at an international level, EXACT

provides for the exchange of reliability test data with a dozen foreign

countries. As our current national policy encourages foreign trade and

the sale of American goods overseas, the EXACT program is one means of

providing greater visibility of our products to potential overseas customers. Therefore, GIDEP is providing test reports to EXACT which document

successful test results on parts and materials. In addition, we are increasing our use of foreign-made parts in our military systems. The availa'

bility of information on those materials is of obvious value. EXACT

provides for data exchange among all member countries, with no increase in

overall cost to GIDEP operations.

The GIDEP program operates as a self-regenerating, closed-loop system.

Engineering, failure, and metrology data are submitted by government or industry participants to the Administration Office for inclusion into the

GIDEP Data Banks. Only data that is germane to the objectives of the

GIDEP program and will be of use to others in the program may be submitted.

Once the data has been screened and indexed, the index is placed in the

GIDEP computer. In addition, the full report is placed in microfilm

cartridges which are distributed along with a hard copy index bi-weekly

to all participants. The Administration Office also distributes a complete

updated index to all participants on an annual basis.

The unique combination of data returned via computerized indexing and

microfilmed reports provides all participants with rapid access to all

GIDEP data. This is accomplished by utilizing their in-house computer

terminal and microfilm reader. Participants may also at their discretion

request a computer index search and copies of various reports directly

from the Administration Office.

The participant may, through a terminal, address any of the four GIDEP

data banks individually or collectively. Further, he may enter a year

date to restrict his data search to relatively current information, or if

he desires, he may also run an entire historical search of all data in

a particular area. The GIDEP computer will automatically search the Data

Banks and if pertinent items are found, the participant will be provided

with an index number and, in the case of an ALERT, a brief abstract of

what the report contains.

The GIDEP computer is programmed so a participant may initiate a search

using any of several approaches: The computer may be addressed to search

by GIDEP generic code, manufacturer's part number, keyword, industry

standard part number, environmental code, or any other fields contained

168
in the data banks. Once information is located, the computer provides

the participant with a microfilm reel and access number in addition to a report number. All participants are provided with microfilm indexes.

Thus, if a participant does not maintain microfilm records, he may request

copies of specific reports or the loan of a microfilm reel from the

Administration Office.

If an index and computer search fails to identify participant-required

data in the GIDEP Data Banks, he may initiate an Urgent Data Request.

One of the most powerful tools of the GIDEP program, the Urgent Data

Request is a means by which a participant may submit an informal request

to all other participants for information on a specific part, component,

material or test instrument. The Urgent Data Request form is initiated

by the participant which he submits to the GIDEP Administration Office.

The GIDEP Administration Office, upon receipt of the Urgent Data Request,

reproduces the form and distributes it to all participants generally

within a period of twenty-four hours.

When a participant, upon receipt of an Urgent Data Request, finds

information pertaining to the particular problem addressed, he forwards

the information directly to the requester. When data are received in

response to an Urgent Data Request, the requester forwards a copy to the

Administration Office for inclusion in the GIDEP Data Banks.

The GIDEP program is much more than technical information and data

banks. It is people cooperating for the benefit of all. One of the key

people in the GIDEP program is the GIDEP Representative. Assigned by

each new participant from its in-house staff, his responsibility is to

determine who in his organization can use and will generate GIDEP data.

He, more than any other person in the program, can directly influence his

company's success or failure within the scope of the GIDEP program. He

must continually promote the system and inform his associates of its

values and advantages. As a GIDEP Representative he also acts as liaison

among his in-house departments and other GIDEP Representatives and the

GIDEP Data Banks. The GIDEP Representative is also the key to documenting

the cost effectiveness of the program. An effective GIDEP Representative

is synonymous with an efficient GIDEP program.

The GIDEP program provides participants with the vehicle to maintain

and improve the reliability of their product and simultaneously minimize

research time and eliminate duplicate testing of parts, components and

materials. The GIDEP program, properly implemented and utilized, produces

impressive cost effectiveness. In 1966 when the program was relatively

new, the documented cost savings to participants, both government and

industry, was $5 million. By 1970, the cost savings had risen to over

$7 million, and last year the GIDEP program had a cost effectiveness of over $10 million.

169
Increased cost savings is only maintained through participant cooperation and efficiency. Examples of past cost savings are numerous and well

documented. One typical case history concerns an engineer who in 1972 was using the GIDEP Data Banks to obtain failure analysis data on a particular microcircuit . The GIDEP analysis data indicated the microcircuit

exhibited excessive leakage but the problem had been corrected by replacing one of the materials used in its construction. Aware that this

particular microcircuit was currently being used by another company for

production of Minuteman ground support equipment, the engineer recognized

a potential problem and issued a GIDEP ALERT. The ALERT was picked up

by the company using the microcircuit, and, as a result, issued a change

order for some 200 systems still in production. Thus, the repair cost

of 200 systems was avoided, at an estimated cost savings to the company

of over $1,300,000. In addition, if these parts had gotten into the

field, the replacement costs would have been astronomical.

Literally hundreds of major cost avoidance cases are documented each

year by participants in the GIDEP program. As more industry and government activities become involved with GIDEP, the effectiveness and efficiency of the program increase. Organizations involved with the Mechanical

Failures Prevention Group may participate in GIDEP. To apply for participation, direct a letter of request to the GIDEP Administration Office.

The letter should be signed by an official duly authorized to commit the

company or agency to the obligations associated with participant status.

Requests for GIDEP program information and participation requirements

should be directed to:

Officer in Charge (Code 862)

Fleet Missile Systems Analysis and Evaluation Group Annex

Attn: GIDEP Administration Office

Corona, California 91720 Telephone: (714) 736-4677

Autovon: 933-4677

170
MECHANICAL FAILURES PREVENTION GROUP

19th MEETING

The Role of Cavitation in Mechanical Failures

National Bureau of Standards, Boulder, Colorado

October 51, November 1 and 2, 1973

W. F. Adler

Bell Aerospace Company

Division of Textron

Buffalo, New York 1U225

E. L. Armstrong

Mobil Research and Development

Corporation

Billingsport Road

Paulsboro, New Jersey O8062

Bruce C. Baird

The Boeing Company

P. 0. Box 587^7

Houston, Texas 770^8

William Belke

Caterpillar Tractor Company

Research Department Technical

Center

Peoria, Illinois 6l602

Jacques Bonnin

Electricite de France

6 Quai Watier - 78 Chatou

FRANCE

David E. Bowen

Ford Motor Company

lkk-25 Sheldon Road

Plymouth, Michigan U8I7O

H. C. Burnett

National Bureau of Standards

Institute for Materials Research

Washington, D. C. 2023^

Robert York Chapman

Acoustical Research and Development,

Code 78O

Naval Submarine Base New London

Groton, Connecticut 063*+0

John Chung

Aerojet General Corporation

SES Division

P. 0. Box 2173

Tacoma, Washington 98^+01

J. David Cogdell

Caterpillar Tractor Company

Technical Center Building E

Peoria, Illinois 6l602

A. F. Conn

Hydronautics, Inc.

Laurel, Maryland 20810

James F. Cor so Naval Ship Engineering Center

SEC 6153

3700 East -West Highway

Hyattsville, Maryland 20782

M. J. Devine

Naval Air Development Center

Warminster, Pennsylvania 1897^-

Stanley W. Doroff

Office of Naval Research

Code I+38

Arlington, Virginia 22217

171
J . L . Duda

Pennsylvania State University

University Park, Pennsylvania 16802

Robert Eilers

Naval Systems Division

Technical Research Group

The Boeine Company

Seattle, Washington 9812*4-

G. 0. Ellis

Vickers Division

Sperry Rand Corporation

Troy, Michigan U808U

Paul Fleming

National Bureau of Standards

Institute for Materials Research

Washington, D. C. 2023^

Kenneth Foulke

Naval Air Development Center

Warminster, Pennsylvania 1897^-

B. Genetti

General Electric Company

2901 East Lake Road

Erie, Pennsylvania 165OI

K. Gentner

Naval Air Development Center

Warminster, Pennsylvania 1897^

L. M. Gillin

Embassy of Australia

1601 Massachusetts Avenue, N. W.

Washington, D. C. 20036

W. A. Glaeser

Battelle Memorial Institute

Columbus, Ohio i+3201

G. C. Gould

General Electric Company

Schenectady, New York 12^01

J. V. Hackworth

Bell Aerospace Company

Division of Textron

Buffalo, New York lk2k-0

J. A. Halat

Vickers Division

Sperry Rand Corporation

Troy, Michigan ^8084

Edward G. Halleck

Fluor Engineers and Constructors

2500 South Atlantic Boulevard

Los Angeles, California 900^0

F. G. Hammitt

University of Michigan

Ann Arbor, Michigan U8105

Henry Hegner

IIT Research Institute

10 West 35th Street

Chicago, Illinois 60616

Anton H. Hehn o General American Transportation

Corporation

71+1+9 North Natchez Avenue

Niles, Illinois 606U8

F . J . Heymann

Westinghouse Electric Corporation

Lester, Pennsylvania 19113

Robert J. Huebner

Caterpillar Tractor Company

72l6 North Crabapple Court

Peoria, Illinois 6l6lU

David G. Jans son Office of Naval Research

800 North Quincy Street

Arlington, Virginia 22219

Robert L. Johnson

National Aeronautics and Space

Administration

Lewis Research Center

Cleveland, Ohio UU135

Frank C. Jones

K West

9371 Kramer Avenue

Westminster, California 92683

172
E. E. Klaus

The Pennsylvania State University

133 Chemical Engineering Building

University Park, Pennsylvania 16802

E. R. Lamson

Naval Air Development Center

Code 3021

Warminster, Pennsylvania 1897^

Jerre F. Lauterbach

Cummins Engine Company

Columbus, Indiana U72OI

Frederick F. Ling

Rensselaer Polytechnic Institute

Mechanics Division

Troy, New York 12181

Harry E. Hunt

Worthington Corporation

hOl Worthington Avenue

Harrison, New Jersey 079^5

Melvin M. Makara

AVSCOM

12th and Spruce Streets

Saint Louis, Missouri 63IO9

Robert E. Maringer

Battelle, Columbus Laboratories

505 King Avenue

Columbus, Ohio i+3201

James William Mason

Hydrofoil Special Trials Unit

Pier 3, Building 580, PSNS

Bremerton, Washington 9831^-

Ralph A. Matzelle

Caterpillar Tractor Company

Peoria, Illinois 6l6lk

J. I. McCool

SKF Industries, Inc.

King of Prussia, Pennsylvania 19^06

Richard S. Miller

Office of Naval Research

Arlington, Virginia 22217

Raymond Misialek

Naval Ship Engineering Center

Philadelphia Division Naval Base

Philadelphia, Pennsylvania 19112

J . E . Newhart

Naval Air Propulsion Test Center

Trenton, New Jersey 08628

Chester R. Oberg

U. S. Atomic Energy Commission

2^-3 R-l, Route 1

Boyds, Maryland 20720

William O'Keefe

Power Magazine

1221 6th Avenue

New York, New York 10020

Elio Passaglia

National Bureau of Standards

Institute for Materials Research

Washington, D. C. 2023^

F. B. Peterson

Naval Ship Research and Development

Center

Bethesda, Maryland 2001^

M. L 0 Picklesimer

National Bureau of Standards

Institute for Materials Research

Washington, D. C. 202 3^

Carolyn M. Preece

State University of New York

Materials Science Department

Stony Brook, New York II790

Edwin T. Richards

U. S. Naval Fleet Missile Systems

Corona, California 91720

C. N 0 Rowe

Mobil Research and Development

Corporation

Box 1025

Princeton, New Jersey 085^-0

173
Souren Sadjian

Frankford Arsenal

Philadelphia, Pennsylvania 19111

Georges Sapy

Electricite de France

17 Avenue du General de Gaulle

927UO Clamart

FRANCE

William T. Sawyer

Holly Harbor

Oxford , Maryland 2165U

Marvin H. Schlichtmann

John Deere Waterloo Tractor Works

Box 270

Waterloo, Iowa 5O7OU

George F. Schmitt, Jr.

Air Force Materials Laboratory

AFML (MBE)

Wright -Patterson AFB, Ohio

R. F. Scott, Jr.

Travelers Insurance Company

\h0\ Elm Street

Dallas, Texas 75228

Cecil P. Shaffer

Army Aviation Systems Command

Box 209

Saint Louis, Missouri 63166

T. R. Shives

National Bureau of Standards

Institute for Materials Research

Washington, D. C. 2023U

Jesse E. Stern

NASA Goddard Space Flight Center

Greenbelt, Maryland 20771

T. E. Tallian

SKF Industries, Inc.

King of Prussia, Pennsylvania 19^+06

G. Jo Tatnall

Naval Air Development Center

Warminster, Pennsylvania 1897^

A 0 Thiruvengadam

The Catholic University of America

Washington, D. Co 20017

Robb M. Thomson

National Bureau of Standards

Institute for Materials Research

Washington, D. C. 2023^

John M. Thorp

U. S. Army Aviation Systems Command

P. 0 o Box 209, Main Office

Saint Louis, Missouri 63166

L. L. Ting

Ford Motor Company

20000 Rotunda Drive

Dearborn, Michigan J+8121

Robert Y. Ting

Naval Research Laboratory

Washington, D. C. 20375

J. Paul Tullis

Colorado State University

Engineering Research Center

Fort Collins, Colorado 80521

David Wagner

Pratt and Whitney Aircraft

East Hartford, Connecticut 06108

J. A. Walowit

Mechanical Technology, Inc.

Latham, New York 12110

Horst Weinhold

TRW

2765I Lake Shore Boulevard

Euclid, Ohio UU132

17^
Robert M. Whittier

Endevco

801 South Arroyo Parkway

Pasadena, California 9H09

M. Wildmann

Ampex Corporation

Redwood City, California 9ho6k

W. A. Willard

National Bureau of Standards

Institute for Materials Research

Washington, D. C. 2023^+

NBS-114A (REV. 7-73)

U.S. DEPT. of comm. i, pijHl.K ATION OR REPORT NO. 2. Gov't Access ioi BIBLIOGRAPHIC DATA No. SHEET I N35 SP-394

4. mil; and subtfi i.i The Role of Cavitation in Mechanical Failures J s. Publication Date

3. Recipient's Accession N<

Proceedings of the 19th Meeting of the Mechanical Failures

Prevention Group, National Bureau of Standards, Boulder,

Colorado, October 31 and November 1 and 2, 1Q73

April 1974

6. Performing Organization C ode

Editors T 0 R. Shives and W 0 A. Willard

Performing Organ. Report Nc

9. PER FORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

10. Project /Task /Work Unit N< 312C400

11. Contract /Grant No. ONR NAonr 70-7^

NASA. S-55581A

12. Sponsoring Organization Name and Complete Address ("Street, City, State, ZIP)

Office of Naval Research, Arlington, VA 22217

NASA Goddard Space Flight Center, Greenbelt, MD 20771

NBS Institute for Materials Research, Washington, DC 2023^

13. Type of Report & Period Covered Final

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES

Library of Congress Catalog Card Number: 74-7057.

16. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here.) These proceedings consist of a group of sixteen submitted papers and discussions

from the 19th meeting of the Mechanical Failures Prevention Group which was held

at the National Bureau of Standards on October 31 and November 1 and 2, 1973 „ The central theme of the proceedings is the role of cavitation in mechanical

failures

.

17. KF.Y WORDS (six to twelve entries; alphabetical order; capitalize only the first letter of the first key word unless a proper name; separated by semicolons)

Bubble collapse; cavitation; cavitation damage; cavitation erosion prevention;

erosion; surface roughness.

18. AVAILABILITY pC Unlimited

[ For Official Distribution. Do Not Release to NTIS

|7 _ Order From "Sup. of Doc, U.S. Government Print ing Off ice Washington, D.C. 20 i02, SD Cat. No. CI 3. 10:3 94

I

' Order From National Technical Information Service (NTIS)

Springfield, Virginia 221^1

19. SECURITY CLASS

(THIS REPORT)

UNCLASSIFIED

21. NO. OF PAGES

183

20. SECURITY CLASS

(THIS PAGE)

UNCLASSIFIED

22. Price $2.25

USCOMM.DC 29042-P74

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