” 

Gun Tube Coatings in Distress 

by Paul J. Conroy, Paul Weinacht, 

and Michael J. Nusca 

AKL-‘l&K-2393 February 2001 

” 

i 

20010430 150 

Approved fw.public release; distribution is unlimited. 

r/ 
(I 

. 

, 

The findings in this report are not to be construed as an official 

Department of the Army position unless so designated by other 

authorizedocumms. 

Citation of mmufwturer’s or trade names does not constitute an 

official endorsement or approval of the use thereof. 

Destroy &is report when it is no longer needed. Do not return 

it to the orlghtor. 
Army Research Laboratory 

Aberdeen Proving Ground, MD 21005-5066 

ARL-TR-2393 February 2001 

Gun Tube Coatings in Distress 

Paul J. Conroy, Paul Weinacht, and Michael J. Nusca 

Weapons and Materials Research Directorate, ARL 

Akmved for Dubhc release: distribution is unlimited. 
Abstract 

. 

The effects of thermal, chemical, and mechanical stress components of erosion on the 

degradation and failure of current and potential future gun tube coating materials are 

investigated. Potential failure mechanisms at the interface are shown and discussed. The 

interfacial coating/base material stress calculation methodology, including coating crack 

formation and coating adhesion criteria, is presented. Possible erosion mitigation 

concepts are also presented to counteracthe formation of cracks in the coating. 

‘ 
. 

Acknowledgments 

The authors would like to thank Dr. Jonathan Montgomery of the U.S. Army Research 

Laboratory (ARL), Aberdeen Proving Ground, MD, for providing material data and insight. 

. . . 

111 
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. 

. 

iv 
Table of Contents 

Acknowledgments ............................................... 

ListofFigures .................................................. 

ListofTables ................................................... 

1. Introduction .................................................... 

2. Structural Calculation Methodology ................................. 

3. Erosion Calculations at Coating Crack Base .......................... 

4. Liner-Substrate Interfacial Degradation Concepts ..................... 

5. Summary ...................................................... 

6. References ..................................................... 

DistributionList ................................................. 

Report Documentation Page ....................................... 

Page 

. . . 

ill 

vii 

ix 

1 

3 

11 

14 

17 

19 

23 

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vi 
List of Figures 

. 

Figure 

1. 

2. 

3. 

4. 

5. 

6. 

7. 

8. 

: 

Thick-Walled Cylinder Radial and Shear Stress Plotted Radially With Constant 

500~MPa Inner Wall Pressure, No Temperature Effects . . . . . . . . . . . . . . . . . . . . 

Maximum Computed Interfacial Shear Stresses for O.OlO-in Chromium and 

Tantalum Coatings With UT’S Values . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 

Description of Postmortem Erosion Pit . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 

Analytical Surface Coating Description Including Stress . . . . . . . . . . . . . . . . . . . . 

Single-Shot Pit Erosion Depth for O.OlO-in Chromium- and Tantalum-Plated 

Gun Tubes Firing 120~mm APFSDS Rounds . . . . . . . . . . . . . . . . . . . . . . . . . . . . 

Substrate Erosion Beneath “Older” Cracks Showing Lateral Distribution . . . . . . . 

Substrate Erosion Beneath “Newer” Cracks Showing Lateral Distribution . . . . . . 

Inter-facial Temperatures for Continuous and Discontinuous Regions With 

DescriptiveDiagram . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 

8 

10 

11 

12 

13 

13 

15 

16 

vii 
INTENTIONALLY LEFT BLANK. 

. . . 

vlll 
Table 

List of Tables 

1. Computed Nominal Mhimum Distance Between Cracks Not Considering 

DilationEffects . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10 

. 

ix 
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X 
1. Introduction 

The thermochemical modeling description used for the calculations in this report has been 

previously reported (Bundy et al. 1996; Conroy et al. 1997, 1998, 1999a, 1999b). The current 

methodology has the addition of an ablative surface coating, which follows the same methodology 

as described in Weinacht and Conroy (1996). This report describes the potential stresses encountered 

by a contiguous coating where residual stresses from manufacturing have not caused cracking or 

delaminated fi-om the substrate. Erosion calculations are presented for both chromium and tantalum 

coatings at the bottom of cracks, which form from thermal-mechanical cracking. A mechanistic 

description of the undermining and subsequent removal of the coating followed by discussions of the 

results is presented. 

Historically, gun tubes were bare steel in the United States until World War II (WWII), when 

many coatings and types of steel were examined. These included Ta, Cr, MO, W, Co, and stellite 

(Bush et al. 1946). As of August 1942 (Ritchie 1942), the only corporation in the United States to 

successfully plate a gun barrel with chromium was Battelle. Coating and erosion mitigation methods 

during the war, and shortly thereafter, involved primarily electroplating chromium or inserting a 

section of stellite into a .50-caliber machine gun. During the postwar years from 1946 to 1965, the 

literature on gun tube wear and erosion became scarce. Due to issues with weapon systems in 

Vietnam (namely the 175~mm gun as well as machine guns), a renewed interest in gun tube wear and 

erosion emerged around the mid- 1960s. Meetings concerning liners, coatings, and additives took 

place (Ahmed and Picard 1970), and much research (Jones and Breitbart 1959; Frankle and Kruse 

1967; Gehring and Meyer 1968; Grabke 1969; Ho&rein and Thruvengadan 1973) was performed 

on the wear and erosion issue. 

Mitigation efforts advanced from 1946 to the early 1980s with the advent oftalc, waxes, and TiO, 

additives and with mixtures involving each (Brosseau and Ward 1975; Lannon and Ward 198 1; Chen 

and Baldauf 1980; Ward and Brosseau 1980). The work from the mid-1960s through the early 1980s 

was summarized in two trisexvice meetings held at Picatinny Arsenal, NJ, by Jean Paul Picard in 1977 

1 
and 1982 (Picard and Ahmed 1977; Picard et al. 1982). From the mid-1980s through 1990, interest 

in gun tube wear and erosion waned due to fielded systems meeting mission requirements. Usually 

when an erosion issue would arise, as with the original M9 19 round which later became the “6 16W,” 

a team was assembled to solve the problem with additives and propellant formulations (Talley and 

Owczarczak 1991). Currently, other than low-contractile (LC) electrodeposited chromium (Chen 

and Baldauf 1980) and the nitrided M242 25-mm barrel, there are no new fielded coatings or new 

methods of plating or coating a gun tube surface. However, there are over half a dozen ongoing 

projects in the U.S. Army ranging from functionally gradient ceramics to magnetron sputtering for 

coating/lining gun tubes. Despite these ongoing projects, it is still difficult to obtain a coatingniner 

which functions as well as the current methods. 

Why is it so difficult to find or develop a good coating? Coatings mitigate erosion for two basic 

reasons. First, a successful coating is not reactive with the propellant product gases. Second, the 

coating serves to distribute and insulate the thermal load to the gun tube base material, thus reducing 

the effective peak temperature experienced by the gun steel. Typically, when the coating/lining 

material has both of these properties, the mechanical properties are not suitable for gun tubes; 

therefore, they are coatings or liners and usually not a significant structural load-bearing section of 

the tube. Another issue, somewhat separate from thermochemical erosion, is the wear due to the 

projectile passage, which implies that the coating must also be resistant to mechanical wear. Aside 

from these basic requirements, there are many other important details, such as coating adherence to 

the base material and the coefficient of thermal expansivity (CTE), which should be close to that of 

the base material to prevent thermal stress cracking and many other details. One final concern is that 

both the coating and the deposition technique must be cost effective (Montgomery 1999). When put 

together, the requirements for a coating are extremely demanding. 

Describing a coating in an interior ballistic setting requires inclusion of the structural 

considerations. Therefore, the following structural derivation is provided to describe the coating 

crack distribution. 
2. Structural Calculation Methodology 

The coupled thermomechanical stress state of the barrel with a coating is required to determine 

the interfacial loads which contribute to coating breakdown. In order to solve for interfacial shear 

stresses between a coating and base material in an axisymmetric plane stress situation, the principal 

stresses must be computed for both coating and base material. It is important to note that the 

calculations contained within this report assume that the initial state of the tube and coating are 

unstressed. That is, residual stress due to plating and autofrettaging is not present. Inclusion of the 

nominal stress state is possible in future work. Following Timoshenko and Goodier (1970), note the 

following equation of equilibrium, where r, 8, and z denote radial, azimuthal, and axial directions. 

3 + or-% - = 0, 

ar r 

(1) 

with the following Poisson relations for E strains, with c1 as the CTE and v the Poisson coefficient, 

and 

For displacement u, 

&r - aT = qJr -V(q) + q), 

E 

&e - 

aT = 1 E(oe-v(~r + q), 

&Z -aT=L . 

E ( ( oz-v (30 + or) ) 

er=$ee=-14,and EZ=Eo 

ar (5) r 

3 

(2) 

(3) 

(4) 
provide a basis from which to compute the stress and displacement relations, assuming plane stress 

( i.e., o, = 0) reduces the stress components to 

% = E EO+V&,-(l +v)aT , 

l-9 

E 0 r = - & +vEe- (1 +v)aT , 

1-3 r 

and 

oz = 0. 

The axial strain relationship reduces to 

(6) 

(7) 

(8) 

E 2 = UT-s oe+cJr . (9) 

The following differential equations for the radial displacement result by substituting the previous 

stress functions (equations 6 and 7) into the equilibrium differential equation using the constitutive 

relations of equation 5: 

a2u l&l u 

r2 

+---- 

rar g 

= a(1 +v)X, 

ar (10) 

or 

i 1 

a 1 84 = a(l ++?g -v 

& r ar ik ’ 

which, when solved for u, results in 

4 
r 

24 = a(1 +v)i 

s 

ATdr + C,r + 2. 

r r 

a 

Through the relations of equation 5, this, in turn, supplies the displacements zr and ee as 

E, =- u(;:v) ’ ATdr s + ‘(l +V)AT + C, - f 

r a 

(12) 

(13) 

and 

g 0 = ‘(l +‘I rATdr + C s + 2 

r2 a 

1 r2 . (14) 

Using these results in stress relations (equations 7 and 8) produces the radial and hoop stress as 

Or = errATdr + 

s f2 r, 

-:(I -v) 

and 

rATdr E -EaAT + - 

l-3 t q1 +v> +?(I -v) I , 

(15) 

(16) 

where the constants C, and C, are found by applying the following boundary conditions at r = ri and 

o, = -Pi and at r = r,, o, = -PO, resulting in 

c = qu-VI Pi(l -9) 

1 rF(l +v) E(1 +v) (17) 

and 
c2 = g’::[ Pi-p.+Eg.mr][ JY]. (18) 

To compute the stresses for the coating (liner) and base material, the interfacial pressure must be 

known. The displacement compatibility at the interface supplies this interfacial pressure by stating 

that the displacement of the coating, subscripts c, must be equal to that of the base material, 

subscripts b, at the interface (i.e., from equation 12): 

where C1, and C,, are 

Ck = C,,<l -v,> 

r2(1 ‘V,) = Alc*C,, - A2c*Pj, 

and 

c = u-4 Ecac ‘I 

2c ECU -v,) 

I 

Pi-PI+--- 

s 

rATdr 

2 

rI ri 

while CIb and C,, are 

(20) 

Ir 2 2 

rZ ri 

2 

rI -rf I = AZb*C2b - AZb*P,, (21) 

Cl, = 

&Cl -v& P,(l -v3 

..rzZ(l +v& E*(l +Q 

= AI&C,, - AZb*P,, 

and 

(22) 

* 

6 
c = u-v3 Ebab ‘I 2 2 

26 E,U -v& 

PI-P0 +- rz ro rATdr - 

2 

$ Ir rz - rz2 

= A3b*Pz - A3b*P, + A3b*A4b. (23) 

rZ rI 

Substitution of the defined variables into the displacement compatibility relation of 

equation 19, along with the following variables for the integral terms, 

‘I 

AOc = a,(1 +v,)-LlrATdr, 

rz r. 

and 

‘I 

AOb = a,(1 +vb)L/rATdr = 0, 

rz rI 

(24) 

(29 

produces the inter-facial stress, PI, as 

A2c*rz+Pi -A3c*(Pj +A4c) -AOc 

Pz = * (26) 

Thus, substitution of equation 26 into C, and C, enables the calculation of the radial and hoop 

stress. The resultant difference between the hoop stress from the coating and base material results 

in the inter-facial radial shear stress for contiguous materials. This analytical formulation was 

tested against ANSYS for constant inner pressure and temperature loads and against Lame’s 

analytical solutions for constant inner pressure loads, as presented in Figure 1. 

An estimate of the axial interfacial shear stress follows from the computation of the axial strain 

for the coating and base material independently, with no axial stress, from equation 4 as 

7 
lOOO-- 

&lo-- 

3 

5% 000-- 

j 

z 

400-- 

200-- 

O0.018 o.oxJ 0.022 0.024 0.02f3 o.oB3 0.030 

Radial Distance (m) 

Figure 1. Thick-Walled Cylinder Radial and Shear Stress PIotted Radially With Constant 

500~MFa Inner Wall Pressure, No Temperature Effects. 

Substituting and reducing equations 15 and 16 into equation 27 for CJ, and o, produces 

(27) 

where AT is the interfacial temperature increase. The equivalent axial stress associated with this axial 

strain is simply eJZ. The difference between the coating and base material’s equivalent axial stress is 

the axial shear, which must be supported by the interface. 

The possible distance between fractures of the coating can be represented by the balance of the 

shear stress, with the ultimate strength (UTS) of the coating axially as 

8 
(2% 

which reduces to the axial interfracture length of 

while radially, the inter-fracture length is simply 

(30) 

(31) 

Figure 2 shows interfacial shear stresses experienceduring the firing of an armor-piercing, finstabilized, discarding-sabot (APFSDS) round in a 120~mm M256 cannon. Note that for high 

contractile (HC) chromium, the stresses never appear to exceed the ultimate values because HC 

chromium is very brittle. The shear stresses possibly exceed LC chromium’s ultimate stress. 

However, note that LC chromium is more ductile than HC chromium and fails at about 10% strain. 

Tantalum shows much higher values of inter-facial shear due to CTE differences. Tantalum also 

exhibits much lower values of ultimate strength. Tantalum, however, is extremely ductile 

(Montgomery 1999), which makes it highly unlikely to crack. 

The coatings do not show the tendency to readily debond or fail in shear at the interface. If they 

did, then the coating would fail rather quicky and would not be successful, and the guns would not fire 

for hundreds of rounds. The coatings appear to relieve stresses through plastic strain or cracking. The 

minimum crack distances for a 120~mm M256 tank cannon firing an APFSDS round in which the 

barrel is coated with 0.01 in of either chromium or tantalum are presented in Table 1. Dilation effects 

were not taken into account in this determination. The larger distance between radial cracks suggests 

fewer radial cracks in the tube, causing more axial cracking. Dr. Paul Cote of Benet Weapons 

Laboratory (Watervliet Arsenal, Watervliet, NY) discussed some of the cracking that he observed 

between 155~mm and 120~mm systems as well as HC and LC chromium (Cote and 

9 
1700 

1600 

1500 

1400 

c 

it3 

- 1300 

x 

E 

m 1200 

8 

a 1100 

ii! looo 

!m 

800 

700 

I’ ’ . . c Chrome Axial a 

1 TantalumHoop .I‘F 

Chrome HOOP 

HC Chrome, UTS Est. 

2336-3023 MPa 

1580 

896 

Ta, UTS Est., 40% E: 

228-900 MPa 

700 800 900 1000 1100 1200 1300 

Axial Distance from Rear Face of Tube (mm) 

1400 

Figure 2. Maximum Computed Interfacial Shear Stresses for O.OlO-in Chromium and 

Tantalum Coatings With UTS Values (Single APFSDS Round From an M256 

120-mm Cannon). 

Table 1. Computed Nominal Minimum Distance Between Cracks Not Considering Dilation 

Effects 

Watweb (1999). 

Rickard 1999). According to Dr. Cote, there are microcracks caused by residual stresses created 

during the plating process. Macroscopic cracks arising from thermomechanical stresses also exist. 

The computed cracking is more of the macroscopic type of cracks found by Dr. Cote than the 

microcracks. These chromium-coating crack results are similar to those observed in gun tubes 

10 
(Sopok et al. 1988). Usually the dimensions of the coating “platelets” are similar to the coating 

thickness. These calculations were performed using nominal physical values which may not represent 

those of the actual as-deposited materials. Given all the intrinsic qualifications of these coating vs. 

coating calculations, tantalum should show less propensity than chromium for cracking. The fact that 

there are cracks in existing chromium coatings and mechanisms to produce these cracks leads to the 

following section in which the computation of erosion at the bottom of the cracks is performed. 

3. Erosion Calculations at Coating Crack Base 

Figure 3 represents what occurs beneath the coating near a crack. A jet of combustion gases 

washes through the crack, subsequently reacting and eroding the substrate material and leaving a pit 

with horizontal extensions near the interface. Many species of materials in the eroded region have 

been identified and include sulfur, oxygen, and talc, which are from the propellant additives (Cote and 

Rickard 1999; Fisher et al. 1982). Sul&r, which penetrates deeply into the eroded pits, comes from 

the black powder ignitors and potassium sulfide flash suppressants used in the charges. I 

Gun Steel 

White Layer Untempered Martinsite 

(Solute of C in FCC Fe with coating of 

Post Mortem Contents: Fe, 0, Talc 

Heat Effected Zone 

‘Fe,C) 

Figure3. Description ofPostmortem Erosion Pit (Source: Dr. Paul Cote, Benet Laboratories). 

11 
Figure 4 is a cartoon of the numerical simulation of the erosion at the bottom of a crack. It is 

important to realize that crack orientation is critical. If the crack is longitudinal, it would enable 

propellant product gases to wash directly in and out with tremendous amounts of heat transfer and 

the possibility for surface reaction, in comparison to a radial crack which would rely more on 

recirculation and diffusion of gases to the interface. The radial orientation is chosen for the 

following calculations to better represent those cracks which are very narrow in nature. All of the 

items in Figure 4 are included in the calculations as well as reactions which occur within the shown 

control volume. The reactants and products for the calculations consisted of H,, 0,, CO, CO,, H,O, 

SO,, N,, NO,Cr,O,, Fe&, FeO(s), FeO(l), Fe,O,(s), Fe,O,(s), Fe, Fe(l), FeO(g), FeO,, H,, NO,, and 

NI(l). Other possible products, such as sulfur-containing species, were not included because the 

interior ballistic calculation of the charge did not contain sulfur fi-om the ignitor. The evidence 

produced by Dr. Cote (Cote and fickard 1999) warrants further investigation of sulfur. 

:_ ,.,: _.__ 1,’ ‘,’ 

I:~:,:s:~:~.,~,.,. ( T&r Stress I: 1 11. 

Figure 4. Analytical Surface Coating Description Including Stress. 

The results of the erosion calculations are presented in Figure 5 and show a general trend to . 

decrease forward from the forcing cone region. The amount of material removed from the bottom 

of a crack is approximately l-l .5 p per shot. This is an “order ofmagnitude” given the firing history 

of 120~nun cannons and subsequent micrographs of subsurface erosion, as shown in Figure 6 and 

12 

. 
1.6 

1.0 

0.8 

600 700 800 !wo loo0 1100 l2aI 1300 1400 

AxialDistanehmRearFaceofTube(mm) 

Figure 5. Single-Shot Pit Erosion Depth for 0.010~in Chromium- and Tantalum-Plated Gun 

Tubes Firing 120~mm APFSDS Rounds. 

Original 500X 

Figure 6. Substrate Erosion Beneath “Older” Cracks Showing Lateral Distribution (Source: 

Dr. Paul Cote, Benet Laboratories). 

13 
provided by Dr. Cote. Figure 5 also shows that due to thermophysical properties between the 

coatings, tantalum might have more rapid deterioration at the bottom of a crack than chromium. 

This confounds the previous tube service life prediction that tantalum tends to crack less than 

chromium, if at all due to its dilation properties. 

The erosion did not predominantly progress downward from the bottom of the crack in the 

micrograph in Figure 6 but rather traversed along the interfacial boundary, essentially debonding the 

substrate from the coating. The crack shown may be an older one which may have been formed due 

to residual stresses from manufacturing. Also important to note from this figure is that the cracks 

can be a small fracture (narrow) in comparison to the coating thickness. The white layer exhibited 

in this micrograph indicates high levels of heat input at the substrate surface, 

Figure 7 reveals significant pit erosion, possibly produced when a crack first forms. Given more 

firings on this particular specimen, the sample may ultimately appear as Figure 6 appears. Jmportant 

to note in Figure 7 is that the interface has been degraded significantly with minimal in-depth erosion 

in the regions away from the erosion intrusion. This crack, which produced the substrate erosion in 

Figure 7, is like Figure 3 where it was filled with material-primarily oxides, iron, sulfides, and talc 

from the charge additive for wear reduction. Iron sulfide melts at 1,463 K, whereas iron carbide 

melts at 1,423 K. This indicates that iron sulfide is certainly an important factor in the erosion 

process and has not been included in these calculations. 

4. Liner-Substrate Interfacial Degradation Concepts 

The degradation of the interface is evident in both Figures 6 and 7. The chemical reactions 

between the propellant products and the substrate gun steel in this inter-facial region are primarily 

caused by the temperature of the substrate. The subsurface temperature has both convective and 

conductive contributions, as in equation 32. 

14 
Figure 7. 

Original 500X 

Substrate Erosion Beneath “Newer” Cracks Showing Lateral Distribution (Source: 

Dr. Paul Cote, Benet Laboratories). 

Q = h(T,,-TJ - KS - 

K a2T 

--. 

r ar2 (32) 

Without a crack or defect present in a 0.01 -in chromium-lined 120~mm M256 tank cannon firing 

an APFSDS round, the peak interfacial temperature is about 800 K, as shown in Figure 8. If the 

substrate material is separated from the coating by a gap, then one of the heat transport mechanisms 

(conduction across the interface) is interrupted. Thus, only indirect and highly inefficient 

conduction, as well as the convection, can contribute resulting in a much reduced substrate peak 

temperature below the region of interruption in contact. If, however, the substrate material is in 

intimate contact with the coating near the degraded interface, then conduction around the edge takes 

place. Note that if the coating is separated, as in Figure 8, then the coating peak temperature rises 

dramatically. An important factor to keep in mind is that this differential in peak temperatures 

between the substrate and coating is influenced by the size of the undermined region. The larger the 

undermined region, the higher the temperature difference is between the coating material and 

substrate. Thus, as the service life increases, the rate at which the undermining occurs will probably 

increase as well. 

At least three hypotheses can describe how the chromium is subsequently removed once the 

coating is sufficiently undermined. The first is the obvious one in which the coating lifts or falls out 

15 
‘//A \ 

Time (ms) 

Figure 8. Interfacial Temperatures for Continuous and Discontinuous Regions With 

Descriptive Diagram. 

due to complete separation from the substrate. The second removal method is more probable, novel, 

and violent. If the cracks surrounding the coating platelet are small enough to enable gas choking 

to occur during blowdown, the orifice under the platelet would act as a mini-bore evacuator and 

retain high pressure. This would expel the remaining platelet off the surface in a potentially dramatic 

fashion during blowdown. Depending upon the ultimate strengths of each, either the substrate or 

coating would fail in tensile fracture or the interface would fail. The ultimate strength of chromium 

is about 413 MPa (Matweb 1999), while that of gun steel is about 1,500-1,800 MPa 

(Halt et al. 1975). This would suggest that bits of chromium might be found on the surface after the 

platelet is removed if it was not completely undermined. The third method is that during the 

corrosion process, the developing and expanding oxide layer slowly pushes the chromium upward, 

enabling projectile engagement. 

16 
5. Summary 

The principal finding and discussion from this research are that neither chromium nor tantalum 

coatings (0.0 10 in thick) fail in axial or radial shear. The coatings crack or plastically deform to relieve 

stress. The computed platelet size is consistent with experimental observations. Tantalum coatings 

appear to crack less than chromium coatings. (Hopefully, they will not crack at all.) The computed 

substrate erosion is consistent with experimental observations. Reaction sets (kinetics) will assist in 

more accurately describing the surface conditions. Sulfur may be a significant player in the surface 

reaction process by producing FeS, which melts at 1,463 K. The coating substrate interface is the 

most likely region of erosion and subsequent failure due to energy transport mechanisms. A literal 

undermining or unzipping ofthe interface occurs through thermodynamically driven thermochemistry. 

Once the coating is undermined, it possibly acts as a pressure vessel afler the passage of the 

rarefaction wave during the blowdown process. This may lead to forced removal of the platelets. 

Subsequent corrosion may push the platelets from the surrounding coating and enable the projectile 

passage to remove it. In this line of reasoning, storage conditions may play some role in postfiring 

oxidation of subsurfacexposed steel, and possibly, a postfiring treatment of the gun can be found to 

control corrosion, which may increase tube life dramatically in the future. 

17 
INTENTIONALLY LEFT BLANK. 

18 
6. References 

, 

Ahmed, I., and J. P. Picard. “Gun Tube Erosion and Control.” Proceedings of the Interservice 

Technical Meeting, Watervliet, NY, 1970. 

Brosseau, T. L., and J. R. Ward. “Reduction of Heat Transfer to Gun Barrels Using Wear-Reducing 

Additives.” Transactions of the ASME, pp. 610-614, 1975. 

Bundy, M. L., P. J. Conroy, and J. L. Kennedy. “Experimental and Simulated Bore Surface 

Temperatures for 120-mm Ammunition.” Journal of Defense Science, vol. 46, no. 4, 

pp. 223-232, October 1996. 

Bush, V., J. B. Conant, and L. H. Adams. “Hypervelocity Guns and the Control of Gun Erosion.” 

Summary Technical Report of Division 1, NDRC Volume l,Washington, DC, 1946. 

Chen, E. S., and W. Baldauf. “Improved LC Chromium for Gun Tube Application.” 

ARLCB-TR-80008, Benet Weapons Laboratory, Watervliet, NY, March 1980. 

Conroy, P. J., P. Weinacht, and M. J. Nusca. “120-mm Gun Tube Erosion Including Surface 

Chemistry Effects.” ARL-TR-1526, U.S. Army Research Laboratory, Aberdeen Proving 

Ground, MD, October 1997. 

Conroy, P. J., P. Weinacht, M. J. Nusca, and K. Rice. “Analytical Thermal and Erosion 

Investigation ofthe Extended Range 5” Navy Gun.” Proceedings of 35th JAANAF Combustioh 

SubcommitteeMeeting, Tucson, AZ, vol. 1, pp. 91-102, 1998. 

Conroy, P. J., P. Weinacht, and M. J. Nusca. “Erosion: A Parametric Study-Flame Temperature.” 

ARL-TR-1954, U.S. Army Research Laboratory, Aberdeen Proving Ground, MD, June 1999a. 

Conroy, P. J., P. Weinacht, and M. J. Nusca. “Erosion Studies of Existing Army Systems (Point 

Studies).” ARL-TR-2054, U.S. Army Research Laboratory, Aberdeen Proving Ground, MD, 

September 1999b. 

Cote, P. J., and C. Rickard. “Gray Layers and the Erosion of Chromium Plated Gun Bore Surfaces.” 

Accepted for publication in JYEXR, August 1999. 

Fisher, R. M., A. Sizrmae, and M. H. Kamdar. “Metallographic Studies of Erosion and Cracking 

of Cannon Tubes.” Proceedings of the Tri-Service Gun Tube Wear and Erosion Symposium, 

Dover, NJ, pp. 151-160,1982. 

Frankle, J. M., and L. R. Kruse. “A Method for Estimating the Service Life of a Gun or Howitzer.” 

Memorandum Report No. 1852, U.S. Army Ballistic Research Laboratories, Aberdeen Proving 

Ground, MD, June 1967. 

19 
Gehring, J. W., and K. H. Meyer. “Study to Increase Gun Barrel Life by Plating the Bore With 

Tungsten.” AFATL-TR-76-124, Eglin Air Force Base, FL, October 1968 @D-851-365). 

Grabke, H. J. “Kinetics of Phase Boundary Reactions Between Gases and Metals.” AGARD 

Conference Proceedings No. 52, France, 1969. 

Ho&rein, A., and A. Thruvengadan. “Development of Scaling Laws for Gun Tube Erosion.” 

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21 
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REPORT DOCUMENTATION PAGE Fom Approved 

OMB No. 0704-0188 

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I. AGENCY USE ONLY (Leave blank) 2. REPORT DATE 3. REPORT TYPE AND DATES COVERED 

February 200 1 Final, October 1998 - October 1999 

4. TITLE AND SUBTITLE 5. FUNDING NUMBERS 

Gun Tube Coatings in Distress 622105AH84 

6. AUTHOR(S) 

Paul J. Conroy, Paul Weinacht, and Michael J. Nusca 

7. PERFORMING ORGANIZATION NAME(S) AND ADDRESS 8. PERFORMING ORGANIZATION 

U.S. Army Research Laboratory REPORT NUMBER 

ATTN: AMSRL-WM-BD ARL-TR-2393 

Aberdeen Proving Ground, MD 210055066 

9. SPONSORING/MONITORING AGENCY NAMES(S) AND ADDRESS 1O..SPONSORING/MONlTORlNG 

AGENCY REPORT NUMBER 

11. SUPPLEMENTARY NOTES 

12a. DlSTRlBUTlONlAVAlLABlLlTY STATEMENT 12b. DISTRIBUTION CODE 

Approved for public release; distribution is tmhrnited. 

13. ABSTRACT(Maximum 200 words) 

The effects of thermal, chemical, and mechanical stress components of erosion on the degradation and failure 01 

w-rent and potential future gun tube coating materials are investigated. Potential failure mechanisms at the interface are 

shown and discussed. The interfacial coating/base material stress calculation methodology, including coating crack 

formation and coating adhesion criteria, is presented. Possible erosion mitigation concepts are also presented tc 

zotmteract the formation of cracks in the coating. 

14. SUBJECT TERMS 15. NUMBER OF PAGES 

xosion, coatings, CTE, surface reaction, gun tube coatings, surface stress, coating failure 34 

16. PRICE CODE 

17. SECURITY CLASSIFICATION 18. SECURITY CLASSIFICATION 19. SECURITY CLASSIFICATION 20. LIMITATION OF ABSTRACT 

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