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Evaluation of Tribological Response of Molybdenum Disulphide-Based Coatings to Varying Environments

Permanent Link: http://ufdc.ufl.edu/UFE0021420/00001

Material Information

Title: Evaluation of Tribological Response of Molybdenum Disulphide-Based Coatings to Varying Environments
Physical Description: 1 online resource (98 p.)
Language: english
Creator: Hamilton, Matthew A
Publisher: University of Florida
Place of Publication: Gainesville, Fla.
Publication Date: 2007

Subjects

Subjects / Keywords: cryogenics, friction, mos2, temperature, tribology, wear
Mechanical and Aerospace Engineering -- Dissertations, Academic -- UF
Genre: Mechanical Engineering thesis, Ph.D.
bibliography   ( marcgt )
theses   ( marcgt )
government publication (state, provincial, terriorial, dependent)   ( marcgt )
born-digital   ( sobekcm )
Electronic Thesis or Dissertation

Notes

Abstract: Molybdenum disulphide-based coatings are the standard solid lubricant coatings for mechanisms intended to operate in low earth orbit. These coatings are known to provide low friction in a high-vacuum environment. In my study, a variety of hard metallic and solid lubricant coatings were tested in a self-mated configuration at temperatures ranging from -80?C to 180?C to evaluate the friction response to varying temperature. The tests were performed using a pin-on-disk tribometer inside an environment chamber at fixed normal load and sliding velocity. Systems where the contact was dominated by plastic deformation and wear debris generation showed no evidence of thermal sensitivity and had friction coefficients ranging from 0.2 to 0.9. However systems where motion was accommodated by interfacial sliding experienced as high as an order of magnitude increase in friction coefficient (0.02 to 0.2) as the temperature of the system was decreased from 180 to -80 degrees C. The low friction coefficients of many solid lubricant coatings have been attributed to interfacial sliding in the material. In these experiments, the goal was to prove the systems were interfacially sliding by analyzing the evolution of surface topography. The results indicated that for the MoS2-based coatings it took an average of 30 cycles to remove a single atomic layer of MoS2 (6 angstroms). This showed that interfacial sliding was occurring in a majority of the contact. The end goal of the study was to comment on the application specific performance of the coatings in a bushing-shaft configuration. In this system, the friction coefficient showed athermal behavior caused by severe wear in the contact area. Ultimately designers of space mechanisms will have to determine the severity of the tribological contact in order to estimate the fluctuations in friction coefficient that may be experienced by varying temperature.
General Note: In the series University of Florida Digital Collections.
General Note: Includes vita.
Bibliography: Includes bibliographical references.
Source of Description: Description based on online resource; title from PDF title page.
Source of Description: This bibliographic record is available under the Creative Commons CC0 public domain dedication. The University of Florida Libraries, as creator of this bibliographic record, has waived all rights to it worldwide under copyright law, including all related and neighboring rights, to the extent allowed by law.
Statement of Responsibility: by Matthew A Hamilton.
Thesis: Thesis (Ph.D.)--University of Florida, 2007.
Local: Adviser: Sawyer, Wallace G.

Record Information

Source Institution: UFRGP
Rights Management: Applicable rights reserved.
Classification: lcc - LD1780 2007
System ID: UFE0021420:00001

Permanent Link: http://ufdc.ufl.edu/UFE0021420/00001

Material Information

Title: Evaluation of Tribological Response of Molybdenum Disulphide-Based Coatings to Varying Environments
Physical Description: 1 online resource (98 p.)
Language: english
Creator: Hamilton, Matthew A
Publisher: University of Florida
Place of Publication: Gainesville, Fla.
Publication Date: 2007

Subjects

Subjects / Keywords: cryogenics, friction, mos2, temperature, tribology, wear
Mechanical and Aerospace Engineering -- Dissertations, Academic -- UF
Genre: Mechanical Engineering thesis, Ph.D.
bibliography   ( marcgt )
theses   ( marcgt )
government publication (state, provincial, terriorial, dependent)   ( marcgt )
born-digital   ( sobekcm )
Electronic Thesis or Dissertation

Notes

Abstract: Molybdenum disulphide-based coatings are the standard solid lubricant coatings for mechanisms intended to operate in low earth orbit. These coatings are known to provide low friction in a high-vacuum environment. In my study, a variety of hard metallic and solid lubricant coatings were tested in a self-mated configuration at temperatures ranging from -80?C to 180?C to evaluate the friction response to varying temperature. The tests were performed using a pin-on-disk tribometer inside an environment chamber at fixed normal load and sliding velocity. Systems where the contact was dominated by plastic deformation and wear debris generation showed no evidence of thermal sensitivity and had friction coefficients ranging from 0.2 to 0.9. However systems where motion was accommodated by interfacial sliding experienced as high as an order of magnitude increase in friction coefficient (0.02 to 0.2) as the temperature of the system was decreased from 180 to -80 degrees C. The low friction coefficients of many solid lubricant coatings have been attributed to interfacial sliding in the material. In these experiments, the goal was to prove the systems were interfacially sliding by analyzing the evolution of surface topography. The results indicated that for the MoS2-based coatings it took an average of 30 cycles to remove a single atomic layer of MoS2 (6 angstroms). This showed that interfacial sliding was occurring in a majority of the contact. The end goal of the study was to comment on the application specific performance of the coatings in a bushing-shaft configuration. In this system, the friction coefficient showed athermal behavior caused by severe wear in the contact area. Ultimately designers of space mechanisms will have to determine the severity of the tribological contact in order to estimate the fluctuations in friction coefficient that may be experienced by varying temperature.
General Note: In the series University of Florida Digital Collections.
General Note: Includes vita.
Bibliography: Includes bibliographical references.
Source of Description: Description based on online resource; title from PDF title page.
Source of Description: This bibliographic record is available under the Creative Commons CC0 public domain dedication. The University of Florida Libraries, as creator of this bibliographic record, has waived all rights to it worldwide under copyright law, including all related and neighboring rights, to the extent allowed by law.
Statement of Responsibility: by Matthew A Hamilton.
Thesis: Thesis (Ph.D.)--University of Florida, 2007.
Local: Adviser: Sawyer, Wallace G.

Record Information

Source Institution: UFRGP
Rights Management: Applicable rights reserved.
Classification: lcc - LD1780 2007
System ID: UFE0021420:00001


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04e623f898740f862fbfc35058c4ab43
2df8e55f1b998cbadc6cb901b6b5edf5b907be08







EVALUATION OF TRIBOLOGICAL RESPONSE OF MOLYDENUM DISULPHIDE-
BASED COATINTGS TO VARYING ENVIRONMENTS






















By

MATTHEW A. HAMILTON


A DISSERTATION PRESENTED TO THE GRADUATE SCHOOL
OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT
OF THE REQUIREMENTS FOR THE DEGREE OF
DOCTOR OF PHILOSOPHY

UNIVERSITY OF FLORIDA


2007































O 2007 Matthew A. Hamilton





































For Hobbes









ACKNOWLEDGMENTS

I would like to thank Dr. W. Gregory Sawyer for his help and guidance throughout this

undertaking; without him, this dissertation would never have been realized. I would also like to

thank my friends and family for their input, encouragement and comic relief. Finally, I thank my

wife Melissa. Her support, above all others, has allowed me to achieve this goal in my life.












TABLE OF CONTENTS


page

ACKNOWLEDGMENT S .............. ...............4.....


LI ST OF T ABLE S ............ ...... .__ ...............7...


LI ST OF FIGURE S .............. ...............8.....


AB S TRAC T ............._. .......... ..............._ 1 1..


CHAPTER


1 INTRODUCTION .............. ...............13....


1.1 Motivation of Research............... ...............1
1.2 Literature Review .............. ...............14....


2 EQUIPMENT DESIGN............... ...............20.


2. 1 Design Philosophy ................. ...............20.......... ....
2.2 High Vacuum Chamber ................. .......... ...............21. ....
2.3 High Vacuum Pin-On-Disk Tribometer .............. ...............23....
2.4 High Vacuum Linear Reciprocator ................. ...............27...............
2.5 Cryogenic Pin-On-Disk Tribometer .............. ...............30............. ...
2.6 In Situ Wear Tribometer ................. ............. .. ...............33....
2.7 Cryogenic High Vacuum Bushing Tribometer ................. ...............36..............

3 VARIABLE ENVIRONMENT EXPERIMENTS .............. ...............44....


3.1 Overview of Materials Used in Experiments ................. ...............44........... .
3.2 Cryogenic Pin-On-Disk Experiment ................. ...............45........... ...
Environmental Protocol s .............. ...............45....

Experimental Procedure .............. ...............47....
3.3 In Situ Wear Experiment .............. ...............49....
Sample Description .............. ...............50....
Environmental Protocol s .............. ...............5 1....

Experimental Procedure .............. ........................5
3.4 Cryogenic High Vacuum Bushing Experiment ................. ...............55........... ..
Environmental Protocol s .............. ...............55....
M material Preparation .............. ...............56....
Experimental Procedure .............. ...............56....

4 RE SULT S .............. ...............58....


4. 1 Cryogenic Pin-On-Disk Experiment ................. ...............58........... ...
Determining a Friction Coefficient .............. ...............58....












Experimental Results ................. ...............58.................
4.2 In Situ Wear Experiment .............. ........................6
4.3 Cryogenic High Vacuum Bushing Experiment ......____ ..... ... ........._.......6
Preliminary Experiment............... ...............6
Solid Lubricant Coating Results ............ ..... ._ ...............69...

5 DI SCU SSION ............... ...............7 1...


6 CONCLU SIONS................ ..............7


7 FUTURE DIRECTIONS .............. ...............80....


APPENDIX


A UNCERTAINTY ANALYSIS OF FALEX POD TRIBOMETER............. .._.........___....82


B HIGH VACUUM PIN-ON-DISK EXPERIMENT ................. ...............85........... ...


Environmental Protocols .............. ...............85....

Experimental Procedure ................. ...............85...
Humid Air and High Vacuum Testing .............. ...............86....
Pump and Purge Test Results ................ ...............86........... ...

C AUGER ELECTRON SPECTROSCOPY OF COATINTGS .............. ...............88....


D GIBBS FREE ENERGY CALCULATIONS .............. ...............91....


LIST OF REFERENCES ................. ...............94................


BIOGRAPHICAL SKETCH .............. ...............98....










LIST OF TABLES


Table page

3-1 List of coatings initially tested for environmental sensitivity .............. .....................4

4-1 All the steady state wear testing in humid air and dry nitrogen. ............. .....................6

5-1 Activation energy values from Monte Carlo simulations of all the data along with the
uncertainties in those values. ............. ...............75.....

B-1 Summary of results from high vacuum pin-on-disk tribometer in humid laboratory air
and high vacuum. All tests were run at a normal load of 5N and a sliding speed of
20 m m /sec. ............. ...............86.....

D-1 List of values for enthalpy and entropy for species in (D-1).(55) ................. ................. .91










LIST OF FIGURES


Figure page

2-1 Vacuum chamber pump down curve with inset of residual gas analysis results at high
vacuum .............. ...............23....

2-2 Pin-on-disk design variations .............. ...............25....

2-3 Components of the high vacuum linear reciprocating tribometer.. ................ ................ ...29

2-4 Schematic of modified pin-on-disk tribometer used in the temperature studies. The
forces in the system are depicted in the image above using grey arrows. .........................31

2-5 Cryogenic pin-on-disk sample housing. ............. ...............32.....

2-6 Overview of in situ tribometer ................. ...............34...............

2-7 Bushing tribometer ................. ...............37................

2-8 Illustration of load path to ground for bushing tribometer design ................. ................ ...39

2-9 Friction coefficient derivation for bushing tribometer results ................. ........... ...........42

3-1 Description of impinging j et technique used to cool and clean the surface of the disk
during experiments............... ..............4

3-2 Screenshot of LabView software written for the in situ wear experiments. ................... .........53

4-1 Friction plot of the electroless nickel coating at -250C. The plot indicates the large
fluctuations in friction coefficient throughout the experiment. The average value for
this data was CI = 0.8 with a standard deviation of 0. 1. The temperature of the system
is also plotted to indicate the small fluctuations in temperature over a 15 minute
period. ............. ...............59.....

4-2 Friction response of metallic coatings to varying temperature. The friction forces were
scattered due to the large amount of surface deformation and wear that occurred
during the test ................. ...............60........... ....

4-3 Friction response of metallic coatings with solid lubricant impregnated. (a) The friction
results for hard anodize with PTFE coating at -300C. (b) The friction results for the
metallic coatings impregnated with PTFE were much more uniform than the metallic
coatings alone, but the coatings did not appear to be sensitive to changing
tem peratures. .............. ...............61....

4-4 Results for bulk polymeric samples against a stainless steel pin. The PTFE-based
samples displayed a linear decrease in friction coefficient with increasing










temperature, while the UHMWPE sample maintained a constant friction coefficient
regardless of temperature. ..........__.......__ ...............62...

4-5 Friction coefficients of MoS2 based solid lubricant coatings at varying temperatures.
The nearly pure MoS2 with nickel coating (95% MoS2, 5% Ni) had the lowest
sensitivity to temperature. The commercially available MoS2 with Sb203 and gold
had the lowest recorded friction at 1800C and the highest increase in friction as
temperature decreased (over an order of magnitude). ................ ................ ........ .63

4-6 Methodology for calculating wear volume. (a) Plot of a line scan taken from the
undeformed surface scan made by the Zygo prior to beginning the experiment and a
line scan taken from the surface scan made after 6,000 cycles. (b) Estimation of the
wear scar cross-sectional area using the difference between the two line scans. This
area is multiplied by the length of the track to estimate a wear volume. ................... ........65

4-7 Plot of volume lost vs. work input into the system. The plot indicates the wear of this
coating is sensitive to the partial pressure of water in the environment. In this case,
the wear rate initially in air is much less severe than transitioning from a dry
environment to a humid one. ............. ...............66.....

4-8 Plot of volume lost vs. work for the MoS2 with Sb203 and gold coating. The uncertainty
in this value is nearly 100% of the value due to the small volume loss over the course
of the entire 2,000 cycles. This result led to testing at higher normal loads and larger
numbers of cycles. ............. ...............67.....

4-9 Initial experiment using fluorinated grease in a high vacuum cryogenic friction
experiment. ........._ ...... .. ...............69....

4-10 Plot of the friction response for all three coatings tested in the experiment. None of the
coatings demonstrated an increase in friction with decreasing temperature in this
geometry. The highlighted outlier in the MoS2 with titanium data set was the last
data point and could indicate the onset of failure for the coating. Many of these
coatings failed after only a few hundred to one-thousand cycles. ............. ...................70

5-1 Illustration of a potential energy barrier typically found in activated chemical processes.
The image to the left of the energy curve is a representation of Boltzmann energy
distribution for a large number of atoms or molecules in a system at different
temperatures. The portion highlighted in black is the area of the curve containing
atoms with energies high enough to overcome the potential energy barrier. ....................72

5-2 Illustration of methodology for creating a Monte Carlo simulation of activation energies. ...73

5-3 One data series in standard and logarithmic plot with fit. ............ ........ .............7

5-4 Activation energy fits for all the MoS2-based composite coatings. This value can be
thought of as a thermal sensitivity for the coating. .............. ...............74....










5-5 Plot of activation energies determined from the cryogenic testing of all the MoS2
coatings as a function of wear rate determined by the in situ experiments for the
sam e coatings. .............. ...............76....

5-6 The deck of cards illustration for sliding between lamellar MoS2 Sheets. The interfaces
that are shearing change throughout the sliding event, but the shear force does not
change. ............. ...............77.....

A-1 Free body diagram of Falex pin-on-disk tribometer with friction coefficient
calculations. ............. ...............82.....

B-1 Pump and purge data results for a thin coating and bulk polymer............_.._._ ........._.._. ...87

C-1 Relative intensities for several elements plotted vs. position across the wear track for the
MoS2 with titanium coating run in humid air environment. Auger data indicates a
high content of titanium and oxygen (likely titania) on the surface inside the wear
track. The signature of the other elements was relatively constant inside and outside
the track ................. ...............88.................

C-2 Relative intensities for several elements plotted vs. position across the wear track for the
MoS2 with titanium coating run in dry nitrogen. Like the humid air data, dry
nitrogen data indicates a higher content of titanium and oxygen (likely titania) on the
surface inside the wear track. However, there are also distinct changes in the sulphur
and molybdenum signatures inside the track indicating more MoS2 is on the surface.....89

C-3 Relative intensities for several elements plotted vs. position across the wear track for the
AFRL chameleon coating run in humid air. There is no signal from the molybdenum
in this coating. Oxygen and carbon are primarily found inside the wear track while
sulphur and antimony are depleted inside the track. ......____ ........_ ................89

C-4 Relative intensities for several elements plotted vs. position across the wear track for the
AFRL chameleon coating run in dry nitrogen. The molybdenum, sulphur and carbon
intensities increase inside the wear track. The difference between this scan and the
open air tests indicate MoS2 is, in fact, drawn to the surface of the coating during
wear ................. ...............90.................









Abstract of Dissertation Presented to the Graduate School
of the University of Florida in Partial Fulfillment of the
Requirements for the Degree of Doctor of Philosophy

EVALUATION OF TRIBOLOGICAL RESPONSE OF MOLYDENUM DISULPHIDE-
BASED COATINTGS TO VARYING ENVIRONMENTS

By

Matthew A. Hamilton

August 2007

Chair: W. Gregory Sawyer
Major: Mechanical Engineering

Molybdenum disulphide-based coatings are the standard solid lubricant coatings for mechanisms

intended to operate in low earth orbit. These coatings are known to provide low friction in a

high-vacuum environment. In my study, a variety of hard metallic and solid lubricant coatings

were tested in a self-mated configuration at temperatures ranging from -800C to 1800C to

evaluate the friction response to varying temperature. The tests were performed using a pin-on-

disk tribometer inside an environment chamber at fixed normal load and sliding velocity.

Systems where the contact was dominated by plastic deformation and wear debris generation

showed no evidence of thermal sensitivity and had friction coefficients ranging from 0.2 to 0.9.

However systems where motion was accommodated by interfacial sliding experienced as high as

an order of magnitude increase in friction coefficient (0.02 to 0.2) as the temperature of the

system was decreased from 1800C to -800C.

The low friction coefficients of many solid lubricant coatings have been attributed to interfacial

sliding in the material. In these experiments, the goal was to prove the systems were interfacially

sliding by analyzing the evolution of surface topography. The results indicated that for the

MoS2-based coatings it took an average of 30 cycles to remove a single atomic layer of MoS2 (6

angstroms). This showed that interfacial sliding was occurring in a majority of the contact.












The end goal of the study was to comment on the application specific performance of the

coatings in a bushing-shaft configuration. In this system, the friction coefficient showed

athermal behavior caused by severe wear in the contact area. Ultimately designers of space

mechanisms will have to determine the severity of the tribological contact in order to estimate

the fluctuations in friction coefficient that may be experienced by varying temperature.









CHAPTER 1
INTTRODUCTION

1.1 Motivation of Research

Designing mechanisms to perform in low earth orbit requires predictable performance

from the sliding interfaces. There are several challenges facing designers of space mechanisms.

First, the temperatures that may be experienced in orbit vary widely (-2000C to 2000C). Second,

the pressures found in low earth orbit are below the limit for many traditional lubricants (1x10-9

Torr). Finally, the mechanisms are typically tested on earth prior to launch, so materials chosen

must not only function in both environments but also resist oxidation and degradation due to

ambient species. Any one of these variables can have an effect on the tribology of a sliding

interface. Characterizing these effects and providing models for predicting performance under

these conditions is the focus of this research.

Many designers characterize frictional losses in a system as a function of the mechanics of

the system. While the mechanics can drive some of these losses, the environment can play an

equally important role in the response of the system. The influence of temperature on the

friction of a system can be dramatic. Polytetrafluoroethylene (PTFE) is a primary example of a

commonly used solid lubricant that is affected by temperature changes. In this polymer,

temperature changes can influence both its mechanical and tribological properties.(1-4)

Molybdenum disulphide (MoS2) has become a standard solid lubricant for high vacuum or

dry conditions. The interest in the tribology of MoS2 can be seen in the large number of

publications focused on this topic. There are however, two areas that have not been exhaustively

studied. The first is the macroscopic friction response ofMoS2 COatings to decreasing

temperatures. The second is an in situ quantification of wear for MoS2 and MoS2-based

composite coatings while cycling environment. A review of current literature yields only one










publication using MoS2 in a CryOgenic environment, but this paper was simply a publication of

test results and offered no hypothesis for the results found.(5) A paper by McCook et al. looked

at PTFE coatings at reduced temperatures and posed the hypothesis that friction of self-mated

PTFE may be thermally activated. The group also notes the appearance of a cutoff temperature

below which the friction no longer rises, but levels out.(6) A more recent work on highly

oriented pyrolytic graphite (HOPG) also noted this type of thermally activated friction response

including a temperature limit, however, this work was on a much smaller scale in ultra-high

vacuum.(7) Another study performed by Schirmeisen et al. found this type of increasing friction

for untreated silicon (111) sliding against a silicon tip.(7, 8) Like the experiment by Zhao et al.

this experiment was performed at ultra-high vacuum (UHV) using an atomic force microscope

(AFM). The results indicate there may be a physical limit where interfacial sliding of these

materials is no longer the most convenient path for energy dissipation. However, it may be just

as possible that the cutoff in each case is caused by separate phenomena. Therefore, the goal of

this study was to determine if interfacial sliding of self-mated materials displayed a thermally

activated behavior and to identify the transition from this behavior to athermal response at the

macroscale.

1.2 Literature Review

Brief History of Molybdenum Disulphide Coatings

As early as 1941 there is reference to the low friction of 'molybdenite' in vacuum. It was

discovered in a search for a replacement lubricant for the rotating anode x-ray tube.(9) One of

the first substantial studies of MoS2 aS a solid lubricant was performed by Johnson, Godfrey and

Bisson in 1948. They looked at the effects of contact pressure and sliding speed on the friction

coefficient for a thin MoS2 COating on steel in air. One conclusion drawn from the study was

MoS2 'is ve7y effective in reducing friction at high sliding velocities. This film material was very









tenacious, was chemically and thermally stable and consequently should have many practical

applications.'(10) After the evaluation of the steady-state tribological properties in air, the

sensitivity of friction coefficient of MoS2 to buildup of surface oxide layers and crystallite

orientation was reported by Johnson and Vaughn in 1956. This study held MoS2 pinS stationary

for varying periods of time and observed the starting friction coefficient and steady-state friction

coefficient. They determined that the 'buildup effect' caused an increase in the friction

coefficient for a short time (less than 10 minutes) before it returned to a steady-state value. Once

the system had run to steady state in one direction, the motor spindle was reversed to analyze the

effect of crystalline orientation on friction. It was determined that orientation had no effect on

steady-state friction.(11) The popularity of the new material led to expansion of its testing.

Deacon and Goodman examined the role of elevated temperature on the friction of MoS2 and

several other lamellar solids. A pin-on-disk style of tribometer was used to measure the friction

of a 1/") diameter hemisphere of platinum sliding on a flat disk of platinum both surfaces covered

with a thin MoS2 layer. The flat disk was heated from room temperature up to 6000C and

friction coefficient was initially above 0.2 but quickly ran down to a value of 0.07 for

temperatures from 400C to 3000C at which point it steadily rose to a value of 0.5 at 5300C. The

results indicated MoS2 COuld maintain a friction coefficient less than 0.1 up to 3000C where the

molecule began to oxidize to MoO3 in air.(12) Because of the distinct difference in the friction

of MoS2 in air COmpared with high-vacuum Haltner and Oliver looked at the effects of the partial

pressure of water on the friction response of MoS2. Pellets of MoS2 were run against a copper

substrate to create a transfer film in air and friction response was observed while varying the

humidity surrounding the sample from ambient levels 25% RH to 0.5% RH. This study posed a

model for friction coefficients as a function of increasing partial pressure of H20 in the









system.(13) Further environmental studies were done with MoS2 in ultra-high vacuum (p < 10-9

Torr) where the temperature was elevated to over 9300C before thermal decomposition was

reached.(14) Ross and Sussman found the oxidation of MoS2 to MoO3 to be a self-limiting

process; meaning the formation of a monolayer of oxide prevented further oxidation. The study

showed steep initial increase in the percentage of MoO3 when exposed to humid air, but as the

exposure time increased, there was no increase in MoO3.(15)

After the discovery of the low friction and thermal stability of MoS2 f11ms it became a

primary candidate for many aerospace applications, particularly elevated temperature systems

such as j et engines. A study of the friction of the available MoS2 COatings at temperatures

ranging from -750C to 2000C was published by Hopkins and Campbell in 1969. A majority of

the coatings in this experiment were MoS2 with graphite using different binders (epoxy, sodium

silicate, etc.). The focus of the study was to determine how different binder materials affected

the friction and wear performance of MoS2 f11ms. The coatings were applied using an air brush

then cured onto the surface. The data from these tests indicated a trend of decreasing kinetic

friction with increasing temperature.(5) Other studies focused on the wear life of burnished

coatings in an effort to improve the life expectancy of the coating.(16) Up to this point, a

maj ority of the studies used manual deposition such as burnishing, painting or evaporating the

material onto the sample. A number of groups were also working on improving deposition

techniques for MoS2. The performance of sputtered MoS2 f11ms was first studied by Talivaldis

Spalvins. His studies showed an increased wear life and a more stable friction response over the

life of the sample for a sputtered MoS2 f11ms in vacuum on a variety of substrates.(1 7, 18) One

important conclusion that appeared in most of the papers on deposited films was cleanliness of

the substrate heavily influenced the performance of the film. Another discovery, documented by









Petrov et al., was the effect of the substrate temperature deposition on the degree of

crystallization and the evolution of the structure inside the fi1m during growth.(19) The study by

Holinski and Gansheimer on the lubricating mechanism of MoS2 indicated the layered structure

shears easily and covers rough asperities on the surface. The interactions between MoS2 layerS

are weak allowing for low shear strength and low friction coefficient. This study illustrates the

sliding as a deck of cards where multiple layers participate in the shearing events simultaneously.

Using SEM, the authors were able to quantify that roughly 25 atomic layers participated in

sliding during the bearings tests performed.(20) Finally, the crystallinity of the MoS2 COating

was interrogated. This study indicated that an amorphous MoS2 had poor tribological properties,

in particular a high friction coefficient (CI = 0.4) and acted as an abrasive that wore the

counterface. Spalvins also studied the wear process for MoS2 f11ms that grew in finger-like

structures from the substrate. The results indicated that these structures break early in the wear

process leaving a well-adhered thin layer of MoS2 bonded to the surface to accommodate the

shear.(21)

The success of new deposition techniques (e.g. radio frequency (RF) magnetron

sputtering(22), DC-diode sputtering(23) and RF diode sputtering(24)) gave researchers the

ability to form composite coatings and study the effects of other constituents added to the MoS2

films. Initially the hope was to reinforce the MoS2 and extend its life, but as new materials were

introduced and deposition techniques were improved another goal arose; to combine other

materials with MoS2 that could help prevent oxidation of MoS2 and provide solid lubrication for

environments where MoS2 WAS not the best choice (namely humid air). This goal has remained a

focus for many groups developing high performance solid lubricant films over the past 30 years.

Zabinski et al. studied the morphological and tribological properties of MoS2 f11ms codeposited









with a variety of materials to identify changes that accompanied improved friction and wear

characteristics. In this study, gold, iron, nickel and antimony trioxide were used as dopants for

MoS2 f11ms. They concluded a decrease in crystalline grain size decreased the friction

coefficient and improved wear resistance of the coatings. In the study a composite of

MoS2/Sb203/Au resulted in the smallest grain sizes of all the coatings.(25) Further studies under

increased temperatures led to the development of a yttria-stabilized zirconia (YSZ), gold, MoS2,

diamond-like carbon (DLC) composite termed a 'chameleon' coating. This coating was

developed at the Air Force Research Labs (AFRL) and produced a friction coefficient of 0. 10 in

air reportedly due to the graphitic carbon from DLC and 0.02 in a dry nitrogen environment

because of MoS2. As the temperature of the system was elevated to 5000C, sliding was

accommodated by drawing gold to the surface yielding a friction coefficient of 0. 15.(26) Many

other studies have been performed using MoS2 in COmbination with other constituents in an

attempt to further improve performance over an increasing range of operating temperatures and

environments.(2 7-31)

Other developments in coating technologies are the use of composites to control the

nanocrystalline structure of the coatings. The AFRL has performed a number of studies on the

structure of coatings using different recipes and correlated that with tribological performance of

the films.(26, 32-34) Another improvement in thin film coatings is customizing the surface prior

to applying the coatings. By designing the surface topography, it is possible to control how the

surface will wear and ensure that the contact is never starved of lubricant. One technique that

has gained popularity is texturing the surface prior to applying coatings. Using a laser, micron-

sized dimples are created on the surface of a solid created. These holes act as reservoirs for the

solid lubricant so it is replenished as the surface wears. This technique, developed by Andre









Voevodin at AFRL, has had success at maintaining low friction for extended numbers of

cycles.(35)



Although the use of deposition techniques has added diversity to the materials chosen to be

added with MoS2 for new coatings, it has also been shown to negatively influence the quality of

the MoS2. Several studies have been performed on testing the quality of MoS2 in deposited

coatings and found that the background pressure of the deposition chamber drastically affects the

ratio of Sulphur to Molybdenum in the coatings. For deposited coatings where the chamber

pressure is on the order of lx10-4 to lx10-3 Torr, the ratio for S to Mo in the resulting coatings

ranges from 1.5 to 1.9. These coatings demonstrated a range of friction coefficient values from

0.007 to 0.1 in inert environments.(36) However, in systems where deposition occurs under

ultra-high vacuum (1x10-9 Torr or less) the ratio is 1.97.(37) The very pure MoS2 COatings had

recorded friction values below 0.002 in vacuum.(36) The poor quality of the sputtered MoS2

leads to Oxygen substitution in the lattice structure and edge sites throughout the coating that are

primary nucleation points for oxidation to MoO3.(38)









CHAPTER 2
EQUIPlVENT DESIGN

2.1 Design Philosophy

The goal in the design of the high vacuum tribometers was to create devices that

functioned inside a vacuum chamber, but required no chamber in particular. In other words, the

tribometers could not require mechanical feedthroughs. There were two reasons driving this

decision initially. The first was that only one vacuum chamber was being fabricated, and to keep

that chamber as simple as possible, no mechanical feedthroughs could be required. The second

was that keeping the motor for the system inside the chamber as near to the sample as possible

minimizing immeasurable losses that arise when complex loading and motion systems are used

to accommodate these feedthroughs. Driving the systems with vacuum compatible motors

simplified the designs guaranteeing that all loads reacted by the sample are read by the load cell

in each system. This design decision requires a motor and stage system that can reliably oscillate

for millions of cycles without adjustment. All the other components can be adjusted prior to

evacuating the chamber; however, a motor failure ruins the experiment.

The philosophy of connecting a load cell directly in the path between the sample and

ground was put forth by Schmitz et al. as one that minimizes the uncertainties in the force

measurements.(39, 40) This philosophy is followed in all the custom-built tribometers discussed

here. To maximize the sensitivity of this tribometer, the load cell is mounted so the friction force

is measured by the most sensitive axes of the load cell. In some cases, the load cell is oriented so

the normal force is also measured using the more sensitive axes, particularly in cases where the

designed normal load is small compared to the capacity of the load cell.

During the construction of the tribometers, the software written for each one was equipped

with the ability to record phase-locked data. This means the positional data is recorded









simultaneously with normal load and friction force. This data is useful for debugging issues with

the design and can expose structure in the data that is caused by synchronous topography in the

sample.

2.2 High Vacuum Chamber

The ambient pressures at 20 km and 2,000 km are 10 Torr and lx10~1 Torr, respectively.

In a vacuum system, pressures above lx10-3 Torr can be maintained through a controlled leaking

of gas into the system. Ultra-high vacuum levels (< 10-9 Torr) are typically obtained through

metal sealed stainless steel chambers that are baked to remove water and continuously pumped.

A decision to operate at high vacuum levels (< 10-6 Torr) was made in effort to minimize

complexity and maximize sample throughput. Because of gas surface interactions (41) that can

often cause dramatic changes in the frictional behavior, the vacuum level or level of cover gas

cleanliness required to simulate space environments continues to be discussed and is likely a

function of sample material, geometry, and sliding speed (some authors suggest that pressures as

low as 10-3 Torr may be sufficient for certain materials (32, 42)).

Controlling the environment around the tribometer is the most critical aspect of vacuum

tribology equipment. For this tribometer a custom-built stainless steel vacuum chamber was

designed to reach the base pressure of lx10-6 Torr in less than two hours. The transition from

viscous flow to molecular flow (occurring ~ lx10-3 Torr) requires a pump for each regime. A

dry-scroll pump was selected for the viscous flow regime (low vacuum) because of its high

pumping speed and it oil free operation, which ensures there can be no back streaming of oil that

would contaminate the chamber.(43) A cryogenic vacuum pump is used after the crossover to

the molecular flow regime (high-vacuum). This system works using a liquid helium compressor

to cool a series of surfaces inside the pump from 77 K to 4 K. The fundamental theory behind

cryogenic vacuum pumps is that desorption of chemisorbed and physisorbed molecules on a









surface is exponentially dependent on temperature(44). There are a number of reasons that this

type of high vacuum pump was selected. First with a turbomolecular pump the system can be

destroyed by a sudden increase in pressure. These pressure rises could be caused by a leak or a

valve being opened mistakenly. The increase in pressure on one side of the system of turbine

blades causes them to deflect enough that they will collide with the second level of blades

resulting in catastrophic failure. However, if this occurs in a cryogenic pump, the pump will be

saturated and can no longer pump gases. The solution to this is to empty or regenerate the pump

by allowing it to warm to room temperature again. The second reason is the pump system works

using cryogenic stages so water is pumped at a much higher rate than with a turbomolecular

pump. Although the pumping speeds for air and nitrogen are equivalent between similarly sized

cryogenic and turbomolecular pumps (1500 L/sec), the pumping speed for water is 266% higher

for the cryogenic pump (4000 L/sec) than it is for the turbo pump (1500 L/sec). In stainless steel

vacuum chambers, water is often the primary source of background pressure for the system. The

pump-down curve for the vacuum system is plotted in Figure 2-1, along with the species found

in the chamber as reported by the residual gas analyzer. Clearly the partial pressure of water is

still the largest contributor of all the species found at this pressure. In an effort to drive as much

adsorbed water as possible from the interior surfaces, heater jackets were wrapped around the

external surface of the vacuum chamber; this helps to helps to drive water off the surface during

pump-down and helps to reduce the rate of water adsorption when the chamber is vented. The

background pressure is read by a pressure gauge mounted directly to the chamber, which outputs

a OV to 10V signal indicating the pressure reading. The system can also determine ambient

species present inside the chamber by using a residual gas analyzer (RGA). The RGA works by

ionizing a sample of the molecules present in the chamber and measuring their mass to charge



















H20
S35


X I I INitrogen (N 14 & N2 28)
~ 17111 IOxygen (O 16 & 02 32)
I II ICarbon Dioxide (COz 44)


H2N | N2 02 CO2
1 5 10 15 20 25 30 35 40 45 50 55 60 65 70
Mass /Charge


ratio. This ratio is often unique to a single molecule or group of molecules and offers insight

into the molecules present due to the environment and those coming from the sample itself.


1000 -

100-

10 -

1 -

0.1 -

0.01-

1E-3 .

1E-4-

1E-5-

1E-6-

1 E-7-

1E-8-


High Vacuum
Pump
(Cryo-Pump)


Viscous Flow
Molecular Flow


Base Pressure
8 x10-8 Torr


1 10 100

Time (min)

Figure 2-1 Vacuum chamber pump down curve with inset of residual gas analysis results at high
vacuum .



2.3 High Vacuum Pin-On-Disk Tribometer

Mimicking a space environment inside the laboratory presents multiple challenges, one of

the most debated issues is the vacuum level required to accurately reproduce space performance.

There is no industry standard for this value, but the Harris Corporation indicated that a vacuum

level of 10-6 Torr is suitable for performance testing of mechanisms. With this in mind, a pin-on-

disk tribometer was designed and constructed that would be capable of operating in this pressure


Roughing Pump
(Scroll Pump)










regime. Many of the friction studies performed on solid lubricant coatings are run using pin-on-

disk tribometers. There are several advantages for this type of tribometer. The disk is only

expected to run in one direction meaning a simple motor can be used. The friction forces are

only measured in one direction so it simplifies the load sensing requirements, and these systems

can typically generate large numbers of cycles in a short period of time. This can be beneficial

when trying to run a low-wear film to failure.

As mentioned each tribometer is driven by a high vacuum compatible motor capable of

operating at pressures below 10-6 Torr. The motor that was chosen was a T-Max 5 servo motor

from Nutec Components Inc. (www.nutecl.com). The unit is a motor and stage combined into a

single housing. The drive is from a linear motor that has been wrapped around a central axis

giving it the ability to turn the stage. This motor is capable of generating 2.9 N-m of torque

continuously. This value is much higher than any torque expected from a pin-on-disk type

contact under loads of 1 to 10 N. The spindle can be commanded to spin at speeds ranging from

0.001 rpm to 1,000 rpm allowing for a wide range of operating speeds. Nutec also outfits all its

motors with positional encoders so the controller can determine the current position at all times.

The encoder built into this system has 72,000 lines resulting in a resolution of 18 arc-seconds.

The load cell for this tribometer is an AMTI (www.amtiweb.com) MC-2.5A six-axis

transducer. The choice to use this load cell was based on availability. At this time no other

manufacturer could guarantee vacuum compatibility to the required levels. The transducer was

modified to ensure vacuum compatibility below 10-7 Torr. The main changes are removing any

paint and decals from the device and replacing all the wiring inside with Teflon coated copper

wires to reduce outgassing. This load cell has a maximum capacity of 450 N in the axial










direction and 225 N in the radial directions. The resolution of the load cell in these directions is

15 mN meaning the lowest resolvable friction reading under a 1 N normal load is 0.015.

To manipulate the sample, the entire assembly is mounted to a set of Schneeberger

(www. schneeberger.com) micrometer stages. The purpose of these stages is to accurately adjust

the track diameter to fix the sliding speed, and to load the sample against the surface after

properly positioning the pin.



(a) noaigsae(b) Pin Sample Disk Holder
6-Channel
Load Cell
a 4 Vacuum Compatible
oo Motor
,*9*o6-Channel
-Load Flexure o Load Cell

X-YTable
Sample Flexure X-Y Table 'o*
Disk Sample- --






Figure 2-2 Pin-on-disk design variations (a) Initial pin-on-disk design used flexures and a
compression spring to apply the load to the system. Also, note the orientation of the
load cell. In this configuration its least sensitive axis is in the normal loading
direction. (b) The final design uses a cantilevered pin holder to apply the load
drastically increasing the stiffness of the system. The load cell is now mounted so the
normal force and friction force are measured by the most sensitive axes of the
transducer.


The machined parts in the pin-on-disk assembly went through a maj or revision. Initially,

the device was loaded using a flexure and compression spring assembly (Figure 2-2a). After

construction it was found that the dynamics of this system made it difficult to accurately measure

friction, particularly if the surface had a misalignment. Any misalignment between the surface

of the disk and the pin resulted in a sinusoidal fluctuation in normal force. The stiffness of the

system was low enough that the vibrations induced by this forcing function caused the pin to









bounce off the disk surface. As a result, a new approach was taken in the revision where a

cantilevered sample holder was designed and used to apply a load to the disk (Figure 3b). In the

new design, the load cell orientation was altered so the two most sensitive axes were measuring

normal load and friction force. The new revision simultaneously simplified the design, improved

the sensitivity of the measurements and increased the stiffness of the system. While any

misalignment between the motor and the pin still caused sinusoidal fluctuations in the normal

load, the value was accurately measured and there was no evidence that the pin ever left the

surface of the disk as it had with the first design.

The sample holders for this tribometer are capable of housing multiple disk sizes as well as

multiple pin sizes. The modular disk housing has a base component that bolts to the face of the

motor. A second component is screwed down over the disk pinching it against the base plate and

aligns the disk face with the face of the motor. The housing can accommodate samples of

varying thickness because of a compressible o-ring of Viton behind the sample that keeps it

pressed to the alignment surface on the housing. The ball holder is machined into the load

flexure for this tribometer. As with the disk holders, there are multiple flexures to accommodate

different diameter pin samples. The sample is inserted into the holder and a set screw is

tightened behind the ball holding it firmly in the housing and preventing it from slipping during

the experiment.

The acquisition of data is done using a 6036E data acquisition card from National

Instruments (www.ni.com). Each of the channels (normal force, friction force, ambient pressure

and motor position) are recorded at up to 10 kHz (a value of 1 k
length of time over which data is collected is variable to allow for different cycle times. A

general guideline was to collect data for a minimum of two cycles to help reduce the effects of









noise in the data as well as any surface anomalies that may skew the result. This data is then

averaged and appended to the average data fie along with a timestamp. Phase-locked data is

also available from this tribometer. This data is periodically collected and stored in a separate

data file for further analysis.

2.4 High Vacuum Linear Reciprocator

A primary issue with pin-on-disk testing is contact pressure for this type of a sphere on flat

geometry is very high. Usually the contact pressures are well above the yield strength of the

coating materials. If a larger pin is used to reduce contact pressures, the sliding speed across the

contact is varying as a function of radius. To counteract this effect, the wear tracks are usually

narrow and can be difficult to analyze. Many microscopy techniques require the area of interest

to be between 100 Clm and several mm for an accurate sampling (e.g. x-ray photoelectron

spectroscopy).

One alternative tribometer design is a linear reciprocator. This type of tribometer allows

for a wide range of pin geometries. These pin geometries can greatly reduce the contact stresses

on the fi1ms and allow a much larger wear track that is more easily analyzed in an XPS or other

spectroscopy equipment. Adjusting the surface temperature of a sample on the linear

reciprocator is also much easier than a pin-on-disk. The stage moves in a linear reciprocation

over a Eixed distance, so flexible cryogenic lines can be attached to the stage for adjusting the

surface temperature. In a pin-on-disk contact, the disk sample is rotating and creating a viable

conduction path for cooling the surface is not as straightforward.

The stage used in this design is a Lineax 10 from Nutec Inc. (http://www.nutecl.com).

The motor and stage are capable of operating at high vacuum levels (pressure < 10-6 Torr). The

motor itself is a brushless linear motor and is equipped with a non-contacting encoder system.

The maximum speed of the motor is 3m/sec with a positional resolution of 0.5 Cm and a










repeatability of 2.5 Cm. The stage is capable of handling a normal force of 1000 N and tangential

force of 325 N. However, as the load is increased, so is the current required to drive the stage,

resulting in larger amounts of heat generated due to resistive heating in the motor. A closed-

circuit cooling system was embedded in the motor housing which can be hooked up to a chilled

water system to maintain a safe operating temperature. This is in contrast to the pin-on-disk

tribometer which does not require cooling due to the low forces and torques generated in those

experiments. The maximum track length of 50 mm is defined by the motor travel. The

mounting plate on the stage has an array of tapped 4-40 holes to allow for a variety of

counterface positions.

The load cell used in this system is the same type as described earlier for the high vacuum

pin-on-disk tribometer, a MC-2.5A from AMTI. The only difference is the load capacity and

resolution of the load cell. This system is designed to work with loads as high as 1000 N, so a

2200 N load cell is used. With all of the load cells from AMTI (the MC-2.5A included) the

resolution in the normal or z direction is twice that of the x and y directions. To enable small

friction coefficients to be measured by this system the more sensitive x and y axes are used to

sense the friction force. As mentioned, the resolution of these load cells is scaled based on their

maximum capacity. The transducer used here cannot resolve a friction force less than 75 mN.

However, the normal loads intended for the tribometer are between 100 and 1000 N resulting in

a minimum resolvable friction coefficient of 0.001 at 100 N.

Due to the size of the linear stage, several loading options were precluded. Ultimately, the

choice was made to use a cantilevered arm loaded with a compression spring, which can provide

a variety of loads by varying the stiffness of the compression spring. The system works by using

a t-shaped arm that pivots about a pair of bearings mounted to the frame of the tribometer. As










the spring is compressed, it imposes a force on the top of the t-shaped component in the

horizontal direction. The body pivots about the bearing axis and imposes a vertical force on the

stage to balance the moment. There is no mechanical advantage in this system; theoretically, the

load applied by the spring is the same as the one applied to the stage. This was done to make it

easier to estimate the required spring stiffness and compression for a desired load. One

advantage of using a simple design is low number of components. Every moving part and

surface interface in a vacuum system is a possible source of virtual leaks and outgassing. By

minimizing the number of mating components, it is possible to improve pump down times and

sample throughput.




6-channel load ce compression spring

sample mutn
plate I QgPP- I I tribometer frame


micrometer stage
sample clamps *:


spring force




cryogenic coolingsyem
normal load


Figure 2-3 Components of the high vacuum linear reciprocating tribometer. (a) Schematic of the
high vacuum linear tribometer. (b) Pin sample holder assembly. (c) The compression
spring pulls the t-shaped lever arm forward imposing a normal force on the surface
between the sample and the counterface.


Although the assembly works using a similar theory to the original pin-on-disk device, the

dynamics of this system should not result in the same issues as the pin-on-disk. This is primarily









due to the higher spring stiffness and large normal loads applied to the surface. Another

difference between the systems is the forcing functions. The pin-on-disk tribometer generates a

sinusoidal forcing function with a frequency equal to the cycle frequency. The linear

reciprocator does not generate the same type of forcing function. Any misalignment between the

surface and the pin will manifest itself as a ramp with a frequency equal to the cycle frequency.

2.5 Cryogenic Pin-On-Disk Tribometer

At the core of the apparatus is a Falex Pin-On-Disk tribometer (see figure 2-5). This

system uses a dead-weight load to apply a normal force onto a pin sample. The pin remains

stationary while the disk below it spins. The shear force acting at the interface between the pin

and disk results in a friction force on the pin. This force pulls the armature attached to the pin

inducing a strain across the load cell connecting the armature to ground. The force is measured

by the deflection of a strain gauge and fed into the internal electronics of the tribometer to

calculate a friction coefficient. The user inputs the mass of the dead-weight load attached to the

armature. The friction coefficient is calculated based on the ratio of the distance from the load

cell to the gimbal and the pin sample to the gimbal. A more detailed discussion of the friction

calculation and the uncertainty in the measurements from this tribometer can be found in

Appendix A.











rn Specifications
y ~Normal Load Range 0.1 N 10N
Pin oldr s Track Radius 0 mm -25 mm
Sf Spindle Speed 10 rpm 200 rpm
0 Strain Gauge Based
a m aLoad Cell Assembly Disk Diameter 1.0"or 2.0"
Disk Holder Frictionless Gimbal Pin Diameter 0.25"
Analog output voltages for
Dead Weight (m) ps,, cycle count


m-g | Counterweight n 2



F_4 2-p-~m-g
s, F, =
3 3


Figure 2-4 Schematic of modified pin-on-disk tribometer used in the temperature studies. The
forces in the system are depicted in the image above using grey arrows.


There were two maj or changes from the as-received Falex system to better accommodate

the testing. The first change was the disk holder. The second generation holder was created for

easy sample changes. The housing itself is made up of two parts. The lower half is mounted to

the spindle driven by the motor. This portion aligns based on the spindle shaft and is tightened

onto the shaft using a clamping mechanism. The top portion is modular so that multiple sample

sizes can be used. Currently, two sample sizes can be used. The first is a 2.0" diameter disk

with a range of thicknesses from 1/8" to 5/16", which was used for all the testing described here.

The second is a 1.0" diameter disk with the same thickness range. The lower half of the housing

has a reverse-threaded stud coming out of the top. The upper half threads onto this stud and is

tightened using a tool provided by Falex. The holes on the outside of the upper half are used by

the tool to apply more torque for tightening the sample housing.

The second major change was the pin holder. The holder was similar to the original Falex

holder in that it had a hexagonal head to allow multiple tracks to be run on a single ball. The









difference comes in how the ball is attached to the housing. The large temperature fluctuations

in these experiments proved a problem for the original holder because the samples were glued to

the housing. It was necessary to change the design and create a mechanical connection between

the ball and housing. This was accomplished by using a threaded stud that runs down the center

of the housing. The balls used must have a 4-40 threaded hold tapped in them, the stud has 4-40

threads on the lower half and left hand threaded '/-20 threads on the upper half. Once the ball is

threaded onto the lower half, the stud is rotated back into the housing causing it to tighten the

ball further and provide a strong force to keep the ball in place. Once the sample is mounted,

decreasing the temperature causes the ball to be further pulled into the housing because the stud

inside is made of aluminum while the housing is stainless steel.




pin housing
ii 1 stainless steel

u pper shaft th readed
1/4-20 (left hand)
aluminum

:----I --- I I -I I\lower shaft threaded
(a)(b) (c) 4-40
th readed 1 /4" ball
aluminum

Figure 2-5 Cryogenic pin-on-disk sample housing. (a) Thread the ball onto the lower shaft using
a standard counter-clockwise motion until it is tight. (b) Rotate the upper shaft in the
clockwise direction to retract the ball into the housing. The ball will fit snugly into
the mating surface inside the pin housing. (c) Section view of the assembly after the
ball has been properly tightened. The materials for each component are also
highlighted to indicate the system will further pull the ball into the housing as the
temperature of the system is decreased.









2.6 In Situ Wear Tribometer

The ability to determine the life expectancy of a solid lubricant coating is of great

importance to designers. Mechanisms must be designed to operate within a useful lifetime, but it

is difficult to accurately predict that life without some guidance on how the system will fail.

Solid lubricant coatings generally fail due to some type of wear mechanism be it delamination,

abrasive wear, plowing etc. Characterizing the type of wear a solid lubricant coating experiences

and its severity can improve designs where these materials are the limiting factor. To

accomplish this, a tribometer was designed and constructed for the purpose of measuring wear of

a surface in situ.

Wear track topography can vary widely based on geometry of the contact and composition

of the surfaces. In the case of many solid lubricants, wear rates are between lx10-4 and 1x10-s

mm3/Nm. This means it could take several thousand cycles before a nanometer of the surface is

removed. Given the sensitivity required to capture the surface evolution, the decision was made

to use a scanning white light interferometer (SWLI). The Zygo New View 5030 was chosen

because of the quality of its optics, and a feature height resolution on the order of angstroms.

This system is equipped with set of motorized stages capable of adjusting the sample orientation

(roll and pitch) and position (x, y, z) to provide the best surface scans. The difficulty was

designing a tribometer capable of functioning in this limited workspace.

Like the vacuum tribometer, finding a motor and stage system suitable for the intended

environment proved difficult. Initially, a small stage was chosen which was driven by a brushed

DC servo motor. The motor was found to have problems running for extended periods in low

humidity environments. This is most likely due to the graphite brushes, commonly found in

many servo motors, becoming brittle in an environment that lacks moisture. The second version

of the stage was a Parker 401-XR linear stage with a HV172 stepper motor. This system is










capable of sustaining a much larger load than the previous version (200 N compared to 5 N) and

came with a linear encoder for monitoring stage position. Measuring absolute position of the

stage provides better repeatability in the measurement that estimating the position using motor

rotations. More specifically, the coupling between the motor and stage has dead zones that

cannot be accounted for when the position is being estimated by motor revolutions. The stage

and stepper motor combination have a positional repeatability of +5Cm. The repeatability is

important when producing the time-lapse images of the track because any systematic drift of the

stage is readily apparent in the resulting video. The new stage also expanded the range of

possible sliding velocities with a maximum velocity of 50 mm/sec.



()6 axis load cell C

leaf-type flexure ~ SWLI objective

micrometer- Ba~counterface
stage


linear stage
pmn sample

(b) adjust stage down
to apply load







Figure 2-6 Overview of in situ tribometer. (a) Schematic of in situ tribometer. (b) Illustration of
loading mechanism. (c) Drawing of tribometer placed on top of the Zygo stage.


The intended normal load for the tribometer was less than 10 N which required a load cell

that could read friction forces as low as 50 mN. A JR3 50M31A load cell was chosen with a

resolution of 28 mN in the normal load direction and 14 mN in the friction force direction.









Following the design philosophy mentioned earlier, the six-axis load cell was placed directly in

the load path between the pin sample and ground. This is believed to be the best way to

minimize uncertainties in the friction and normal force measurements.

The application of load in this tribometer had to be small to allow the system and a

surrounding environment chamber to fit inside the framework of the SWLI. The size constraints

on this tribometer limited the options for the load application mechanism. The design employs a

leaf-type parallelogram flexure attached to a micrometer stage to apply a load to the pin sample.

One benefit of this type of design is the wide variety of normal loads that it can apply by

changing the geometry of the leaf flexures. A second property of this type of flexure system is

its high stiffness in the friction force direction. This will help to prevent the system from rotating

as the force between the pin and the counterface increases. By adjusting the micrometer stage,

the pin sample is brought into contact with the counterface. The load increases as the deflection

in the leaf flexures increases. The six-channel load cell is used to read the applied normal load in

real-time to allow the user to accurately reset the load after the pin has been brought off the

surface for imaging purposes.

The pin sample holder is a PEEK component machined to hold the ball at a 600 angle with

the surface. The purpose for this is to maximize the number of tracks that can be run on a single

ball. By angling the ball in this way, a single sphere may be used for up to 6 tests. This matches

the number of tracks that can be run on the rectangular coupons. The counterface is mounted to

a machined plate attached to the motorized stage. The hole-pattern on the plate is such that the

counterface can be held in three different positions. Each position can have two tracks run on it

for a total of six tracks per sample.









2.7 Cryogenic High Vacuum Bushing Tribometer

Solid lubricants and fluorinated greases are commonly used in bushing configurations to

provide continuous operation in extreme environments where many traditional lubricants are

unable to operate. Thermal limits (high and low) often preclude the use of many hydrocarbon

oils. High vacuum environments are particularly challenging and there are a number of moving

mechanical assemblies that utilize bushings to provide low torque operation in space and high

altitude vehicles.

Bushing contacts are a typically designed to be closely matched axis-symmetric bodies of

revolution (shaft and through hole). Such a common and practical device, bushings are used in

everything from door hinges to jet-engine actuators. Surprisingly, there is little published on

component level testing of bushings; perhaps, due to the multitude of quiet complexities found in

such simple and ubiquitous components. These complexities include evolving geometry during

operation (45),and uncertainties in contact area, pressure distribution, and frictional forces. In an

effort to study the extended performance of a single bushing component, a cold thermal vacuum

bushing tribometer was constructed.

The design of the tribometer followed the methodology described by Schmitz et al.(39),

which essentially describes the importance of having the load path flow through a 6-channel

load-cell that reacts the normal and frictional forces and moments near the point of contact. A

six-channel load cell suitable for high-vacuum operation was designed and fabricated by AMTI

(Boston, Mass) and having a maximum load of 1000N in the loading direction and maximum of

torque of 50Nm. The load cell is a strain-gauge based six-channel device used to sense the load

that is transferred from the flexures into the contact between the bushing and the shaft. It

measures the normal load applied to the bushing as well as the frictional torque generated at the










bushing-shaft interface. The forces and torques are output by the load cell electronics as analog

voltages that can be recorded as the test is running.

Traditionally, dead weight loads, pneumatics, or hydraulics are selected to apply loads in

tribological testing. The large load requirement for this application (500N) precluded the use of

dead weight loading systems in the vacuum chamber, and the modular methodology described in

the introduction precluded pneumatic and hydraulic loading/feedthroughs in the system. A

relatively simple spring loaded system was selected to apply a normal force to the sample; such

systems are typically avoided because creep and wear in the sample act to reduce the strain in the

spring and the normal force is continuously decaying during operation. In this system the

continuous measurement of normal force reduces these time varying biases. Additionally, very

soft springs with the appropriate load capacities were selected to maximize spring deflections at

load and thus minimize variations in load due to wear and other gradual deformations that occur

in during testing.


Flexible Shaft Coupling
Load Cell

Vacuum Motor
-4 o .- Bushing Housing



Load Spring


Alignment Flexure Load Flexure







Figure 2-7 Bushing tribometer









Due to the challenges associated with maintaining low friction forces in vacuum

environments, efforts were made to eliminate bearings and bushings in the design. The

tribometer uses bearings only in the support of the rotating shaft and the motor. To apply normal

load a series of two flexures act as a pivot to transmit the vertical load and as a restraint to force

linear motion of the bushing in the loading direction. The load spring is compressed by

tightening a nut on a threaded rod that is affixed to the load flexure. The alignment flexure is

designed such that it can only move in the horizontal direction and imparts a purely horizontal

load on the load cell and sample assembly through a point contact with a sphere. The function of

the alignment flexure is to accommodate deflection during the wear of the bushing. The flexures

were wire electro-discharge-machined from bulk monolithic stainless steel pieces.

The drive system for the shaft uses a high-vacuum compatible servo motor capable of

running at speeds as low as 0.001 rpm and as high as 1000 rpm. The motor was originally

designed to be cooled via convection. However, since this is not possible inside the vacuum

system a chilled water circulator and aluminum block at the base of the motor are used to

conduct heat out of the motor at high-vacuum levels. A flexible coupling is used to connect the

motor to the rotating shaft and provides the opportunity to easily vary the diameters of the shafts.

The shaft is aligned using two high vacuum rolling element bearings lubricated with fluorinated

grease. These greases have very high molecular weights (3000+ AMU) with vapor pressures

that are below 7x107 Torr at room temperature.

The entire bushing assembly is bolted directly onto to the front of the load cell to reduce

uncertainties in the friction coefficient measurement. As described by Schmitz et al.(39), multi-

axis load cell should be place directly in the load path between the sample and ground as close to

the contact as possible. An illustration of the load path is shown in figure 2-9. All loads reacted









by the shaft are carried through the load cell, which is supported and constrained to linear motion

by the alignment flexure assembly.


Figure 2-8 Illustration of load path to ground for bushing tribometer design.


There are a number of vacuum tribometers; however, the ability to perform cryogenic

vacuum testing is much more limited. To reach temperatures below -100oC liquid nitrogen was

fed through a cold flask within the vacuum chamber. This flask is mounted approximately 75

mm from the sample and thin copper braids are bolted from the reservoir to a copper bushing

housing. This short conduction path and high thermal conductivities of copper made rapid

cooling (~10oC/min) of the bushing assembly possible. To prevent cooling of the load cell, a

polyetheretherkeytone (PEEK) insert is bolted between the copper bushing housing and the load

cell as an insulator. The temperature is read from a thermocouple rigidly held to the outside of

the bushing. The temperature rise across the bushing can be estimated using a simple 1-D heat









transfer analysis. The value for this temperature rise is 30C using a 100 N normal load, 10

mm/sec sliding velocity and a friction coefficient of 0.5.(46)

Acquiring data during an experiment requires a card capable of reading multiple analog

signals simultaneously. Using data acquisition all the different outputs (8) are read at up to 10

k
the cycle time varies from one test to the next depending on the desired spindle speed. Typical

settings for these options are an acquisition rate of 1,000 Hz per channel for a time period of six

seconds. A maj ority of the tests run have been at 20 rpm, so the data is acquired for two

complete cycles before being processed. All of the data collection is phase locked with the

motor position. Plotting the force and torque values with respect to the motor' s angular

coordinate allows identification of persistent features. Any eccentricity between the shaft and

bushing will manifest itself as a sinusoidal fluctuation in the normal force when plotted versus

the angular position of the shaft. These fluctuations are by design small compared to the applied

normal load.

This data collection scheme occupies a great deal of memory. The approach that is

frequently used computes average cycle values from the k
in a single file. Periodically, the phase locked data is stored in a separate file that is time

stamped with the cycle number. The distance value is calculated by using the input spindle

speed, the shaft radius and the time the test has been running. This value has uncertainty because

the values of spindle speed and shaft radius are assumed instead of monitored throughout the

test, however, distance is not a factor in calculating friction coefficient. Schmitz et al. performed

a rigorous uncertainty analysis on a system using the same electronics as the bushing tribometer









and found the contribution of the uncertainty from the electronics to be negligible when

compared with other error sources.(39)

Computation of Friction Coefficient

The strain-gauge based load cell outputs forces and torques about all three axes, however,

some of this information is not used in the computation of friction coefficient. In this tribometer,

the two force values Fx and Fy are used to calculate the total normal force exerted on the shaft.

The only torque that is of interest is the one generated by the friction between the shaft and

bushing, thus only three of the six load cell channels are read into the software.

The friction coefficient is defined as the dimensionless ratio of the friction force between

two bodies divided by the normal force pressing them together. The normal load exerted on the

shaft by the bushing is measured directly, and the frictional stresses result in a torque. Assuming

both the shaft and pin are rigid bodies (illustrated in the figure 2-10b) the normal load and

friction force are assumed to be point loads and the standard friction coefficient equation, a ratio

of the friction force to the normal force, is given by Equation 1.

Ff T
(2-1)


For two bodies in contact that are not infinitely stiff the contact area is finite due to

deformation. This deformation introduces an error into the friction coefficient value reported by

Eqn. (2-1). A simple model is to assume the shaft contacts the bushing over the bushing' s entire

length and the contact occurs over a specific wrap angle. Without more advanced material

properties and contact mechanics, the exact wrap angle and pressure distribution cannot be

determined. However, assuming a uniform pressure distribution over a contact area an analytical

solution can be derived. For the case where the contact happens from -a to a, the normal load









measured by the load cell is the integrated contribution of the pressure distribution illustrated in

Figure 2-10c.




Fn,


Tn P




w Rs

(a) (b) (c)
Figure 2-9 Friction coefficient derivation for bushing tribometer results.


The normal load measured by the load cell is a function of the length of the bushing (w), the

radius (Rs), the pressure (p) and the contact angle (a).


F = 2 p-w-R4cos (0)d6 (2-2)


The average pressure (p) can be calculated as a function of a given the measured normal load Fn.


p = (2-3)
2 w Rs sin (a)

The frictional shear stresses are simply the product of the friction coefficient (CI') and the normal

pressure (p); thus, the frictional torque can by found from the following integral.


T= 2 ~fdO=(2-4)
S2 w 4sin(a) sin(a)

Following Eqn. (2-1), the computed friction coefficient (CI) is given by Eqn. (2-5).


# = p'. (2-7)
sin(a)








For small wrap angle (a) the error is likely negligible, and the reported friction coefficient (CI) is

always larger than the true value (CI'). A plot of the percent error (defined as

%eurror = 100 (p -r p')l/p) is shown in figure 2-11 illustrating that at a wrap angle of 200 the

error in the reading is still below 5%.


35%-

30%-

25%-

20%-

15%-

10%-


O



O


5%-


200 400 600 80o

contact angle (a)
Figure 2-11 Error in friction coefficient based on the contact angle between the bushing

and shaft.









CHAPTER 3
VARIABLE ENVIRONMENT EXPERIMENTS

3.1 Overview of Materials Used in Experiments

This series of experiments was meant to survey the field of possible coating options to

determine their performance under varying environmental conditions. Groups from Harris

Corporation, the Air Force Research Laboratories (AFRL) and the University of Florida were

gathered to give input into the material selection for these tests. The coatings ranged from hard

metallic coatings used in gears to intricate composite coatings designed specifically for low

friction in varying environments. All coatings were applied to 1/4" spheres and 2" diameter

disks of aluminum 7075. This substrate was chosen because it is common in many space

mechanisms due to its high strength-to-weight ratio. The coatings were applied to both the pin

and counterface to examine how the materials respond in a self-mated contact condition, as

opposed to the coating against a nascent aluminum surface.


Table 3-1 List of coatings initially tested for environmental sensitivity
Group Sample Composition Sample Provider
Hard anodized aluminum Hohman Coatings
Hard metallic coatings Electroless nickel Hohman Coatings
Titanium nitride with electroless nickel Hohman Coatings
Hard anodized aluminum with PTFE Hohman Coatings
Metal coatings with PTFE
Electroless nickel with PTFE Hohman Coatings
MoS2 with Sb203 and graphite Air Force Research Labs
MOS2 with Sb203 and gold Hohman Coatings
MoS2-based coatings MoS2 with Sb203 Air Force Research Labs
MOS2 with nickel Hohman Coatings
Mo52 with titanium TEER Coatings
Bulk PTFE University of Florida
Bulk polymers PTFE with PEEK and MoS2 UniVersity of Florida
UHMWPE McMaster Carr

Other Near friction less carbon Argonne National Labs










Some of the coatings listed are commonly found in space applications, hard anodized

aluminum with PTFE, MoS2 with titanium, MoS2 with Sb203 and gold, etc. The MoS2 with

Sb203 and gold coating was originally developed at the Air Force Research Labs, and has

become a commercial standard for low friction interfaces in a variety of mechanisms. While

many of the composites listed are commercially available they are not currently used in space

applications and some are still in developmental stages (MoS2 with Sb203 and graphite and near

frictionless carbon). The purpose of this variety was to give designers insight into the

performance of these new materials as possible options for future designs.

These coatings are created using a variety of different techniques. The anodized coatings

are created by soaking the sample in a sulphuric acid bath while putting a current through the

system. The coating forms about 2Clm thick, while roughly 1l m of the sample material is

removed leaving the sample oversized by 1l m. The electroless nickel process is a similarly

bathed in solution. The system does not require electricity which reportedly reduces the friction

coefficient over electroplated Watts nickel. Both of these coatings can be created with PTFE as

a solid lubricant on the surface.

3.2 Cryogenic Pin-On-Disk Experiment

Environmental Protocols

Reducing the temperature of any surface at atmospheric pressure (as opposed to vacuum)

can form water on that surface as water vapor in the surrounding environment comes out of

solution and condenses. When performing experiments at reduced temperatures it is important

that the test be run above the dew point for the water vapor in the atmosphere to avoid

confounding the tribological results by forming water at the interface. As the surface

temperatures decrease, the amount of water vapor required for condensation to occur decreases.

To combat this problem, it is necessary to remove as much water from the surrounding









environment as possible. The enclosure used in this experiment was a Vacuum Atmospheres

environment chamber, which isolated the experiment from ambient humidity. This chamber was

backfilled with ultra-high purity nitrogen gas (99.999% pure) as well as boil-off from liquid

nitrogen dewars to remove as much water vapor from the system as possible and create an inert

environment with controlled relative humidity levels around the tribometer. The environment

chamber was also fit with a humidity sensor by GE Sensing capable of resolving relative

humidity less than 0.1%.

Inside the chamber, a technique similar to that used in metal inert gas (MIG) welding was

employed to keep the surface of the disk as cold and clean as possible. Liquid nitrogen flows

from a pressurized dewar through braided steel lines to a nozzle located above the disk surface.

The liquid nitrogen flows from the nozzle two inches above the disk surface nearly parallel to the

plane of the disk. Another j et of nitrogen from an ultra-high purity nitrogen cylinder is pointed

down onto the center of the disk. The flow rate of this jet is controlled by an adjustable flow

controller to increase or decrease the amount of gas impinging on the surface. The two jets

intersect above the disk surface causing the liquid nitrogen to volatilize into cryogenic nitrogen

gas which is carried to the surface by the impinging gas nitrogen j et (see Figure 3-1). Using this

method, surface temperatures could be controlled down to -800C. To increase the temperature of

the surface, a heat gun was mounted above the disk surface impinging directly onto the center of

the disk. The temperature of the disk surface was increased as high as 1800C during the testing.

In both cooling and heating, it was important to keep the gas flow near the center of the disk

because any viscous flow off-axis to the load armature could affect normal force and or friction

force.













Nitrogen


Figure 3-1 Description of impinging j et technique used to cool and clean the surface of the disk
during experiments.


Experimental Procedure

The pin-on-disk experiments conducted on the Falex instrument were run at the lowest

obtainable relative humidity levels (below 1% for all the experiments). A 500 gram dead-weight

load was hung from the armature on the tribometer resulting in a nominal normal load of 2.5 N at

the pin. Each sample was tested at three separate temperatures -700C, -300C and 200C. Fresh

tracks were used for each test to eliminate the possibility previous temperature affecting the

results. The procedure for each new track was to run the sample in at room temperature to wear

through any oxide layers that may have formed on the surface between deposition and

experiment. After the friction coefficient reached a steady-state, the test was stopped and the

temperature of the system was reduced to the target value. The temperature was held near the

target for several minutes prior to beginning the test to ensure both the disk and ball were

equilibrated. Once the system was at the target temperature, the motor was started and friction

coefficient recorded for a period of five minutes or more. This was to ensure there were no

transient behaviors in the friction coefficient and there was enough data minimize the effects of

noise in the system. The sliding speed for the innermost track on the sample was 12.5 mm/sec (5









mm radius at 50 rpm), but the sliding speed increased as the track diameter increased because the

motor was unable to reliably spin the disk at speeds less than 30 rpm. The maximum sliding

speed for the samples at the outermost track was 30 mm/sec (10 mm radius at 30 rpm). Johnson

et al. showed friction coefficients for MoS2 COatings were relatively insensitive to sliding speed

over this range.(10) The steady-state temperature rise in the system can be estimated using a

simple 1-D conduction equation.(46)


AT ~(3-1)


To use this equation, the nominal contact radius and the heat flux must be calculated. The

contact radius (a) can be estimated using a Hertzian contact solution. For a sphere on flat

contact, the equivalent radius of the two bodies is equal to the radius of the sphere.

111 11
=-+- =-+-- (3 -2)
R R R, R o 0

The composite modulus for the bodies is calculated using the modulus and Poisson's ratio for

both bodies.


E' = (3-3)
Eb 1-v)+Ea1-vn

Using these values, the contact radius can be calculated given a normal load.(47)


a = 3FRE i(3-4)


The heat flux generated is a function of the average contact pressure, the friction coefficient and

the sliding velocity.

ij= p-P-v (3-5)









In the given experiments, a worst-case scenario for friction coefficients was on the order of 0.5.

The highest recorded sliding velocity was 30 mm/sec, and the normal load was 2.5 N. Given all

these values, the contact radius and heat flux can be calculated.

a = 0.0515mm (3 -6)

if = 4.5 W/mm2 (3 -7)

The resulting steady-state temperature rise for these contact conditions is less than log

(0.9660C). In the case of the solid lubricant films, the friction coefficients were an order of

magnitude lower than the example; indicating the temperature of the system before motion and

after motion should be equivalent.

After running the experiments at reduced temperatures, the materials that displayed a

strong sensitivity to temperature were tested at an elevated temperature of 1800C using the same

procedure described above. This was done to evaluate how much of a reduction could be seen at

elevated temperatures.

3.3 Inz Situ Wear Experiment

The friction response of MoS2 COatings in dry (RH less than 1%) environments and high

vacuum has been well documented. Since the discovery of this material more than 60 years ago,

different techniques have been developed for applying a MoS2 COating, different additives have

been mixed with MoS2 to improve its friction response and its surface adhesion. However, the

wear analysis of this coating has been very limited. A majority of the studies where wear

volume is quantified report values determined ex situ and usually after the coating has failed.

The purpose of this experiment was to capture the evolution of the wear scar as the test

progressed, and determine a correlation between friction coefficients and wear rates.










The tribometer used in this experiment was a linear reciprocating tribometer mounted

underneath a Zygo New View 5030 scanning white light interferometer. Although the resolution

of the device is on the order of angstroms, the device is capable of detecting surface topography

changes on the order of nanometers. The general idea behind this experiment is to image a

section of the wear track at different cycles throughout the test to estimate a volume loss and a

wear rate.

Sample Description

The counterface samples used were nominally 1.5" x 1.0" x 0. 1875" coupons made of 304

stainless steel. The coupons were polished on a polishing wheel to a surface roughness less than

50 nm prior to being shipped to the suppliers for coating. The pin samples were '/" 6061-T6

aluminum spheres that were drilled and tapped for a 4-40 screw. No surface preparation was

done to the spheres prior to coating. The coatings applied to the pins and coupons were all MoS2

based, but had a variety of other constituents. The commercially available coatings used were

MoS2 with titanium, MoS2 with Sb203 and gold, and MoS2 with nickel. The newly developed

coatings from the AFRL were MoS2 with Sb203 and one of the "chameleon" coatings (MoS2

with Sb203 and graphite). All coatings were nominally 1l m thick, although each one was

deposited using different techniques. The MoS2 with titanium is a layered coating created by

sputtering a layer of pure titanium on the surface of the substrate, then co-depositing a layer of

MoS2 and titanium, then a layer of pure MoS2. The MoS2 with nickel and the MoS2 with Sb203

and gold coatings are sputtered coatings of MoS2 with other constituents to improve adhesion to

the substrate, environmental sensitivity and toughness. The AFRL coatings are created by laser

ablating a target made of the desired constituents resulting in a coating with roughly the same

composition as the target.









Environmental Protocols

There were two environmental protocols used to test these coatings. The first was a

cycling environment similar to the 'pump and purge' experiments described in the high vacuum

pin-on-disk experiments; in these experiments the acrylic chamber surrounding the tribometer

was back-filled with dry nitrogen to a relative humidity less than 2%. The sample was run for a

total of 500 cycles in this environment and surface scans were taken at 1 20, 30, 40, 50, 100,

200, 300, 400 and 500 cycles. The chamber was then purged with laboratory air (RH > 20%)

and the sample was run for another 500 cycles using the same scanning frequency. This process

was repeated once more for a total of 2000 cycles. The hope was the data would provide insight

into how the system wears in the transition between a dry environment, where friction

coefficients were below 0.05, and a humid environment where friction coefficient was above 0.1.

Experiments were run starting in nitrogen or humid air to see if the initial run in affected the

future performance of the coating.

The second procedure was aimed at determining a steady-state wear rate for each coating

in a dry nitrogen environment and a humid air environment. Two separate tracks were run on

each sample. The first track was run in dry nitrogen at a relative humidity of less than 1%. The

second was run in laboratory air at a relative humidity of greater than 20%. Each sample was

run for a total of 10,000 cycles and images were taken every 1000 cycles. The chameleon

coating from the AFRL was the coating of interest in this experiment primarily because it is

hypothesized that the coating draws the favorable solid lubricant to the surface of the fi1m based

on environment. If this theory is correct, the track run in air should have a higher carbon

signature than the one run in nitrogen because the graphite in the material is designed to be the

solid lubricant in the moist environment. Appendix C shows the results of Auger electron

spectroscopy run on the Chameleon coating and the MoS2 with titanium coating.










Experimental Procedure

Before any testing could begin with the coated samples, a calibration step was taken to

ensure the obj ective lens of the Zygo was aligned with the wear track, and to determine the

distance between the Zygo lens and the pin holder. An uncoated counterface and pin were

mounted to the system and the counterface was moved under the Zygo obj ective. The roll and

pitch of the Zygo stage was adjusted to align the counterface normal with the obj ective axis.

After the surface was brought into focus, the stage was returned to its initial position and the pin

was loaded to 5 N. The sample was then run for 50 cycles on a 5 mm track to ensure a

noticeable wear scar was generated. The pin was unloaded and brought up from the sample

surface. The stage was j ogged 0.5 mm at a time until the scar was inside the field of view of the

Zygo lens. The Zygo stage was adjusted in the x-y plane to bring the end of the scratch into the

center of the field of view of the Zygo. The distance between the pin and lens was then

determined and recorded for future use during the experiment.



























Figure 3-2 Screensho o~firrn Lbie Flrsoftwar wrte o h nst ereprmns

Whe caibatin ws ompeteth saplecontefae ws bltd t th mtorzedstge.Th
coaed lumnumpinsamleE waslm screed to ~ th PEE houin and the hosn montdoh
load cel. At his pont, th Lab~iw softare wa opene to beintheeprmn.Isd h
software there n ar evrlstp equrdt ei et

1.Etrteclirto osansfrtelad el temcu les umiiysno n
p~ositonal encoder.nrlb~
2.~~~~ Se h oainfrdt ie ob ave n nerayntsontepnad onefc
"samples ~ I,~ i
3.q~ Tae h la cl









4. Define the absolute motor positions of the Zygo obj ective lens, the track start point and
the track end point so the motor can be commanded back to those positions for the
duration of the experiment

5. Move the sample under the Zygo obj ective and record the cycle 0 image of the track.
This image represents the unworn surface topography and will be used to make all the
differential volume loss calculations.

6. Command the sample back to the cycle start position.

7. Bring the pin sample into contact with the surface and adjust the micrometer stage to
apply the appropriate load (in this case 5N).

8. Define the track length, sliding velocity and approximate cycle time for the test, 5mm,
10mm/sec and 1.0 seconds respectively.

9. Enter the data acquisition parameters (sampling period of 1.1 seconds and sampling
frequency of 1 k
10. Define how often the software saves a complete cycle of positional data as opposed to
simply an average value for a cycle.

11i. Enter the position of the center point of the wear track and the percentage of the wear
track to analyze (this is to avoid using the data at the reversal points in the average
friction calculations)

12. Begin the test.

During these experiments, the LabView software recorded normal force, tangential (friction)

force, ambient temperature, relative humidity and stage position. As mentioned, the software

periodically records the phase-locked data associated with a single cycle. Phase-locked data

refers to a correlation between the values recorded, specifically friction force and normal force,

and the stage position. This can be most useful for initial run in or the onset of failure when a

single cycle can be evaluated to look for a portion of the track that is anomalous. Ultimately the

goal of the instrument is to identify portions of the track with erratic behavior and be sure to

image them to give some insight into the failure mechanisms of the coatings.










Following completion of the test, the samples were removed from the tribometer and vacuum

sealed to help protect the surfaces from contamination and oxidation. Two of the coatings were

evaluated using Auger electron spectroscopy to try and identify the composition of the surfaces

after wear occurred.

3.4 Cryogenic High Vacuum Bushing Experiment

Environmental Protocols

Relating friction results obtained using a tribometer in a laboratory environment to actual

frictional losses in a mechanism is not straightforward. Bridging the gap between contrived

experiments and real world environments is a primary goal for many researchers. To that end,

experiments were run using the cryogenic high vacuum bushing tribometer described in chapter

2.4. The purpose of these tests is to compare results obtained at cryogenic temperatures on a pin-

on-disk tribometer with frictional torques generated in a bushing-shaft contact under similar

conditions.

The simulation of a space environment has a broad definition; in this experiment a high

vacuum environment over a range of temperatures are the target conditions. The vacuum level

of lx10-6 Torr was chosen as the required pressure level prior to starting the test. At this

pressure, the monolayer formation times are on the order of one second, meaning a single layer

of molecules (usually water) will adsorb on any available surfaces in roughly one second. The

vacuum pumps continue to work throughout the experiment and can reach levels below lx10-7

Torr. The point of this pressure range is to ensure that the shaft is not covered with more than a

monolayer of contaminant while it is not in contact.

The temperature profile used in these experiments is a ramp that begins at room

temperature (200C) decreases linearly to -600 and then steadily increases back to room

temperature. Throughout the course of this temperature ramp, data points are taken at 200C,










00C, -200C, -400C, -600C, -400C, -200C, 00C and 200C. Although data is collected the entire

time, thermal drift in the system can only be eliminated using reversal techniques. The purpose

of taking data on the reduction in temperature and the increase in temperature is to ensure that

friction response is due to a temperature effect and not to unrecoverable damage to the coatings.

Material Preparation

The bushings used for these experiments were 440C stainless steel with a 10.0 (+0.05/-

0.00) mm inside diameter. The shafts were made of custom 455 stainless steel; they were 8"

long and had a 10.0 (+0.00/-0.05) mm outer diameter. The nominal clearance between the shaft

and bushing was 0.05 mm in all cases. All the shafts were polished to a surface roughness of

less than 100 nm prior to being sent out for coating. Both the bushings and shafts were coated

with the same material to mimic the self-mated contacts of the pin-on-disk experiments. Only a

subset of the MoS2 COatings were chosen for the bushing experiments mainly to verify the

temperature sensitivity of the coatings in an application specific contact geometry. The most

sensitive coating (MoS2, Sb203 and gold), the least sensitive coating (MoS2 and nickel) and an

intermediate coating (MoS2 and titanium) were used to coat bushings and shafts. All three of the

coatings are commercially available and commonly found in high vacuum environments.

Experimental Procedure

The process followed for the cryogenic high vacuum bushing experiments was outlined by

the Harris Corporation. This protocol was also used in the fluorinated grease experiment

mentioned earlier. The procedure for each test was:

1. Open the high vacuum chamber.

2. Load the shaft into the alignment bearings

3. Couple the shaft to the vacuum motor

4. Attach the bushing housing to the load cell










5. Tare the load cell

6. Insert the bushing into the housing

7. Open the LabView software written for the experiment

8. Enter the calibration constants for the load cell, motor encoder, pressure gauge and
thermocouples

9. Define the location the data files are to be saved

10. Apply the normal load to the bushing

11. Close the chamber

12. Evacuate the chamber to high vacuum

13. At the desired pressure, run 5 revolutions in the forward direction and 5 revolutions in the
reverse direction. The average value determines any misalignment or drift in the torque
cell .

14. Begin the test.

15. At each temperature run the bushing in a clockwise direction to steady-state and then
reverse the motor direction and again run to steady state. This allows for an accurate
assessment of frictional torque and reduces the influence of drift.


The bushing was run to a steady-state friction value at room temperature. Once a friction

coefficient was determined at room temperature, the motor was stopped. The liquid nitrogen was

turned on and the indirect cooling system began to reduce the temperature of the bushing

housing and the bushing itself. A thermocouple placed against the outside edge of the bushing

was used to estimate the temperature in the contact. After the temperature on the thermocouple

reached the desired running temperature, the motor was started. The friction was run into a

steady-state value while the temperature was held relatively constant (fluctuations on the order of

+ 50C). The temperature profile for the system was 200C, -600C, -200C, 00C, 200C.









CHAPTER 4
RESULTS

4.1 Cryogenic Pin-On-Disk Experiment

Determining a Friction Coefficient

The calculation of a friction coefficient value for each temperature was accomplished by

determining the steady-state value of the friction coefficient data over a period of time where the

temperature of the system had reached equilibrium and remained constant. The purpose of this

technique was to eliminate transients in the friction response due to fluctuating temperatures.

The mean and standard deviation of the friction coefficient values were calculated for all the data

collected in this period to quantify the consistency of the friction. For cases where the surface

was plastically deformed and/or large amount of debris were generated, the spread on the data

tended to be large compared to the value itself. The solid lubricants, on the other hand, tended to

have a consistent friction response with small standard deviations given a constant temperature.

Experimental Results

The thermal response of materials is a primary concern for designers of equipment for

space applications mainly because the operating range for some of these mechanisms is -1000C

to 2000C. This large range of operating temperatures causes engineers to consider thermal

effects that may not ordinarily be problematic. One such issue is the mismatch of coefficient of

thermal expansion between dissimilar metals. This mismatch can distort the geometries when

the system experiences temperature fluctuations on the order of several hundred degrees.

Another commonly overlooked property is the friction response of a material as the temperature

fluctuates. Many designers view friction coefficient of a material as a property of the material

and not a value that is strongly influenced by geometry and environment. These experiments

began with an experiment performed at the University of Florida using thin PTFE composite










coatings indicating a sensitivity of friction to temperature.(6) While this result is not widely

accepted, and others have published data to the contrary, it lead to the idea that other materials

might also display such a response.(48)

The first coatings evaluated in this experiment were hard metallic coatings. The

temperature of the samples was decreased from room temperature down to 200K (-730C). An

example of the friction result is plotted in figure 4-1. This is the result for the electroless nickel

coating, and indicates the degree of scatter in the data. Each data point on the plot represents the

average value for multiple cycles of data. These results are representative of all the hard metallic

coatings tested, and there was no indication that temperature affected the friction response.

1.2 25

+-* 1.0-

C -15


.9 0.4
o (o
~Friction -"_ 5
LL0.2
Temperature
0.0 .75
0 150 300 450 600 750

Cycles

Figure 4-1 Friction plot of the electroless nickel coating at -250C. The plot indicates the large
fluctuations in friction coefficient throughout the experiment. The average value for
this data was CI = 0.8 with a standard deviation of 0.1i. The temperature of the system
is also plotted to indicate the small fluctuations in temperature over a 15 minute
period.

The erratic behavior of the friction of each of the metallic coatings coupled with small

changes in the friction coefficient over the range of temperatures tested lead to the decision that

no higher temperature testing was warranted for these coatings. There was no indication that the










plastic deformation or abrasive wear of metallic coatings responded to temperature fluctuations.

While this result is not particularly interesting, it provided an important null result; which was

the tribometer itself had no inherent bias resulting from a varying temperature. This can be seen

from the fact that although the temperature of the system changed by 1000C, the responses from

each of the metallic coatings stayed constant and those values were unique to each coating.

1.2-

1.0 Test Conditions

S0.8 -1Normal Load (Fn) = 2.5 N
Q Sliding Speed (v)= 25-50 mm/sec
U 0.6 -Relative Humidity (#)= 0.5%

0.41 Q Hard Anodize
O TiN with Electroless Nickel
0.2-
A Electroless Nickel
0.0 ......
200 220 240 260 280 300 320
Temperature (K)

Figure 4-2 Friction response of metallic coatings to varying temperature. The friction forces
were scattered due to the large amount of surface deformation and wear that occurred
during the test.

While metallic coatings are commonly used in gear teeth because they are known to be

hard and tough, they are also known to have high friction. One method for improving the

performance of these coatings is to deposit a solid lubricant with the coating. Two of the above-

mentioned metallic coatings are also available with PTFE impregnation. The hard anodize and

electroless nickel coatings are commercially available with PTFE. The manufacturers claim

these coatings are highly wear resistant, and the addition of a solid lubricant greatly improves the

friction performance of the coating. These coatings were tested following the same procedure as

the metallic hard coatings to explore the thermal sensitivity. Like the metallic coatings, there

was little evidence to support a thermal sensitivity in the results from these coatings. There is,





















pr~mrmwa~tom~aha~an~cn~c~ta~


- ~O Hard Anodize with PTFE
[] Electroless Nickel with PTFE


however, evidence to support the addition of PTFE greatly decreases the friction coefficients of

these coatings.


25






-25






-75


(a) 0.30-

C 0.25

0
UE 0.15 _

S0.10


LL0.05-

0.00 --


Friction-)-

Temperature


450


600


300

Cycles


0.40 -


u 0.20

O

'E0.10
00


-80 -60 -40 -20 0 20

Temperature (oC)

Figure 4-3 Friction response of metallic coatings with solid lubricant impregnated. (a) The
friction results for hard anodize with PTFE coating at -300C. (b) The friction results
for the metallic coatings impregnated with PTFE were much more uniform than the
metallic coatings alone, but the coatings did not appear to be sensitive to changing
temperatures.

Bulk polymeric components can also be used in mechanisms designed for high vacuum such as

bushings, snaps and pins. These materials tend to be very inert and have low outgassing rates.

The bulk polymeric materials tested in these experiments were PTFE, UHMWPE and a

PEEK/PTFE. Unlike the coatings tested in these experiments, these samples were run against









stainless steel pins. Although the contact does not start out self-mated, polymeric samples

readily form transfer films on the steel surface so after a short run-in period they are essentially

running in a self-mated configuration.(49)

0.18-
O O PTFE
r 0.15 PTFE/PEEK/MoS2
o.1 UHMWPE

U 0.09-

.9 0.06-

LL 0.03-

0.00 111111
-100 -50 0 50 100 150 200

Temperature (oC)

Figure 4-4 Results for bulk polymeric samples against a stainless steel pin. The PTFE-based
samples displayed a linear decrease in friction coefficient with increasing
temperature, while the UHMWPE sample maintained a constant friction coefficient
regardless of temperature.

The PTFE and PTFE/PEEK/MoS2 COmposites both demonstrated a linear trend of

decreasing friction with increasing temperature. Unfortunately, there are so many properties of

PTFE that change with temperature; it is possible to generate a number of explanations for this

result. The UHMWPE, on the other hand, did not show a discernable trend of friction with

temperature. As with the metallic coatings, this is further proof that the system is not inherently

sensitive to temperature.

In contrast to all the previous results, many of the MoS2 based coatings were very sensitive

to temperature fluctuations. MoS2 with nickel was the least sensitive, and reacted similarly to

the PTFE based composites. The trend for this coating was nearly linear with temperature. The










MoS2 with titanium and the MoS2 with Sb203 both had a friction coefficient of 0.05 at room

temperature and increased by 300% as the temperature dropped to -700C. The MoS2 with Sb203

and graphite had a friction coefficient below 0.03 at room temperature, but increased to over 0. 15

at -700C. In the most extreme case, that ofMoS2 with Sb203 and gold, the friction increased by

an order of magnitude from 0.02 to 0.2 over the 1000C temperature drop.
0.25 M S 0.25 M 2+i0.25 MoS2+Sb2,O

li 0.20 6 0.20 a 0.20
S0.15 E 0.15 r ~0.15
.n 0.10 g. 0.10 0.10

0.05 0.05 0.05
0 0 0
-100 -50 0 50 100 150 200 -100 -50 0 50 100 150 200 -100 -50 0 50 100 150 200
temperature OC temperature OC temperature OC
0.25 0.25
0 ~MoS,+Sb,O,+Gr MoS2+Sb20,+Au
.9 0.20 ~~0.20
8 0.15~ o 0.15 Test Conditions
.n 0.10 -1.o 0.10 Normal Load (Fn) = 2.5 N
0.05 A1 i 0.05 .Sliding Speed (v) = 20 50 mm/sec
o~ a Relative Humidity (())= 0.5%/
-100 -50 0 50 100 150 200 -100 -50 0 50 100 150 200
temperature OC temperature OC

Figure 4-5 Friction coefficients of MoS2 based solid lubricant coatings at varying temperatures.
The nearly pure MoS2 with nickel coating (95% MoS2, 5% Ni) had the lowest
sensitivity to temperature. The commercially available MoS2 with Sb203 and gold
had the lowest recorded friction at 1800C and the highest increase in friction as
temperature decreased (over an order of magnitude).

One observation of note was that the materials with the lowest friction coefficient, and

highest thermal sensitivities did not have noticeable wear scars. This is in contrast to the

metallic coatings that exhibited high friction coefficients, no thermal sensitivity and large

amounts of debris generation. The differences in wear debris and wear track generation sparked

an interest in studying the evolution of the wear scars in situ.









4.21in Situ Wear Experiment

The in situ experiments performed in a changing environment provided some insight into

the surface topography changes of MoS2 based systems that accompanied changes in humidity.

While an increase in humidity immediately triggered an increase in friction, the surface

topography did not react as quickly. In fact, the wear mechanism described by Spalvins where

the columnar growths within the film shear and break leaving a well-adhered, thin layer of

composite seemed to be accurate for the MoS2 and nickel coating.(21) The other composite

coatings where the structure is intentionally amorphous seemed to deform initially under loading

and shear stress, but stabilize quickly to a low wear configuration. Another attribute of the

composite coatings is the presence of additive materials. These other constituents (e.g. titanium

or Sb203) are thought to improve the quality and toughness of the coating and prevent the

columnar growth of the film.

Estimating the wear volume given the topographical information of a section of the wear

scar required a technique initially introduced by Williamson and Hunt, and refined by Sayles. In

the original publication from Williamson and Hunt, the technique was used to evaluate the

persistence of asperities after plastic deformation had occurred.(50) Sayles expanded the

technique to evaluate a surface before and after plastic deformation.(51) The method uses an

initial surface topography scan as the basis for the wear scar and takes subsequent topography

scans to calculate the cross sectional area of the wear scar. Using this area extrapolated over the

length of the scratch, a volume loss can be estimated. This method was followed for all the

coatings tested in both the alternating environment experiments and the steady-state wear rate

testing.

The wear rates reported in this section were obtained using a methodology originally

developed by mathematician Stanislay Ulam for predicting odds for the appearance of various










cards in games of solitaire.(52) The technique was named Monte Carlo simulation for Ulam's

uncle, who was a known gambler. The Monte Carlo simulation style is suited to systems that are

dominated by random events or where no analytical solution can be found. The technique uses

random number generation to predict how a 'random' system will react to different inputs. The

reaction is simulated a large number of time and the statistics of the results are considered the

odds of one result or another. For the wear rate data, each data point on the graph represents an

interrupted measurement. Using an estimation of the uncertainty in this measurement and

random numbers, new data points were generated. For each data set, 1,000 possible sets were

generated and a linear fit was used to calculate a wear rate for each. The average wear rate value

was the wear rate calculated from the original data points, and the uncertainty in that wear rate

was the standard deviation of the 1,000 wear rates calculated from the generated data. This

procedure was followed for all the wear rates calculated in these experiments.

(a) soo -wear scar limits(b


o -i8gpebPBh wear scar area

-500-


-P1000 -**

*,. 4 2 o undeformed
-1500 surface
surface after
6000 cycles
-2000 ...
U 170 340 510 680 850 200 300 400 500 600 700
Position (Clm) Position (plm)

Figure 4-6 Methodology for calculating wear volume. (a) Plot of a line scan taken from the
undeformed surface scan made by the Zygo prior to beginning the experiment and a
line scan taken from the surface scan made after 6,000 cycles. (b) Estimation of the
wear scar cross-sectional area using the difference between the two line scans. This
area is multiplied by the length of the track to estimate a wear volume.










By alternating the environments it was possible to determine if the wear rate of the

coatings reacted to the environment as the friction coefficient obviously changed. In a maj ority

of the coatings, it was found that even though friction coefficient changed dramatically, the wear

rate of the systems was not affected. The only coating that demonstrated a noticeable and

repeatable change in the wear rate with environmental changes was the MoS2/titanium coating.

This coating showed a lower wear rate in dry nitrogen when compared with humid air. The

coating also wore more readily when transitioning from a dry environment to a humid

environment.


8 K = 4 x107 K = undefined K= 2 x 10 K= 3 x 10-'~ Test Conditions
In 6 -1 normal load (Fn) = 2.5 N
S sliding speed (v) = 10 mm/sec
r humidity of dry nitrogen < 2%
humidity of lab air > 20%

E 21 -v~c~ O Raw Data
o- Mean Fit
> 0air dry nitrogen air dry nitrogen
0 20 40 60 80 100
Work (Fn*d)

Figure 4-7 Plot of volume lost vs. work input into the system. The plot indicates the wear of this
coating is sensitive to the partial pressure of water in the environment. In this case,
the wear rate initially in air is much less severe than transitioning from a dry
environment to a humid one.

Many of the other coatings did not show appreciable changes in wear rate in varying

environments. For example, the MoS2 with Sb203 and gold showed no evidence of different

wear regimes throughout the 2,000 cycle test regardless of environment. This sample had a wear

rate of 5x10-s mm3/Nm, but the uncertainty in this value was nearly 100 percent. The deepest

penetration depth on this coating throughout the test was 80nm, and the repeatability of the

instrument is estimated at 10nm. One explanation for the large uncertainty is the methodology

used for estimating wear volume was not sensitive enough to reliably capture surface topography










changes that are on the order of 5nm. To increase the wear volume the number of cycles was

increased to 10,000 in each environment and the normal load was increased to 5N. This allowed

for the calculation of a steady state wear rate in humid air and dry nitrogen to discern if the

environment had any effect.


E K= 5 x108 Test Conditions
Su(K)=4 x 10 normal load (Fn) =2.5 N
S2 n siin seedd i-- 10 mm/sec
humidity of lab air > 20%

O Raw Data
o- Mean Fit
> 0air dry nitrogen air dry nitrogen
0 20 40 60 80 100
Work (Fn*d)

Figure 4-8 Plot of volume lost vs. work for the MoS2 with Sb203 and gold coating. The
uncertainty in this value is nearly 100% of the value due to the small volume loss
over the course of the entire 2,000 cycles. This result led to testing at higher normal
loads and larger numbers of cycles.

The 10,000 cycle experiments were aimed at determining a steady-state wear rate in humid

air and dry nitrogen. The hope was to determine if environment had any effect on the generation

of wear debris. In a graphite system, the lack of humidity causes the graphite to become brittle

and wear more rapidly. It was expected that MoS2 based systems would demonstrate a similar

behavior when exposed to a humid environment. This hypothesis proved to be correct in every

case where MoS2 WAS the only solid lubricant present in the system. The wear rates for all these

coatings were lower in dry nitrogen than it was in humid air. Only the MoS2 with Sb203 and

graphite demonstrated a consistent wear rate regardless of the environment. This coating is also

the only one with a constituent solid lubricant suitable for both environments. The theory that

graphite is drawn to the surface during sliding in humid environments is supported by the Auger

analysis in Appendix C. The signature from MoS2 in the track was much stronger in the dry









nitrogen track than it was in the humid air track. Only carbon and oxygen had significant peaks

in the wear track run in humid air. The results for all the steady-state wear studies are shown

(Table 4-1). The uncertainty values listed are the values from the Monte Carlo simulations;

however, based on the sensitivity of the instrument the minimum uncertainty that should be

expected is 1x10-' mm3/Nm.


Table 4-1 All the steady state wear testing in humid air and dry nitrogen.
Khumid air u(K) Kdry nitrogen u(K)
Coating Material (mm3/Nm) (mm /Nm) (mm /Nm) (mm3/Nm)
Mo52 with nickel 2.51x10-s 1.10x10-6 7.44x10-6 1.10x10-6
MoS2 with 5b203 9.18x10-' 1.10x10-' 5.32x10-' 1.90x10-8
Mo52 with titanium 4.23x10-6 3.20x10-7 7.21x1 0-' 1.00x10-7
MoS2 with Sb203 and graphite 1.75x10-* 2.20x10-' 2.29x10-8 4.40x10-9
MoS2 with 5b203 and gold 3.88x 10-" 1.10x10-8 1.86x10-8 5.50x10-9



4.3 Cryogenic High Vacuum Bushing Experiment

Preliminary Experiment

To demonstrate the capabilities of the cold thermal vacuum bushing tribometer, an

experiment of steel on steel with a vacuum compatible fluorinated grease lubricant was

performed. The normal was 100 N leading to a nominal contact pressure of 1 MPa

( j~(2 Rs -1I)) at a sliding velocity of 10 mm/sec. Friction measurements were taken over a

range of temperatures (200C to -750C) and plotted to reveal the influence of temperature on the

friction coefficient of grease. Reversal techniques are used to eliminate biases in the moment

zero. For each data point on the plot, the computed friction coefficient was taken after a sliding

distance of two meters. The confidence intervals shown on the plot are an indication of the

standard deviation of the phase locked friction coefficient data collected at each temperature.



















.0 -0.45 -
0.55 -L 4'hU"SLO I -IIIC~~
-0.65
0 90 180 270 360

0.45 -Angular Position




g 0.35-



u- 0.25-




0.15 I I I I I I I I.
-120 -80 -40 0 40 80
Temperature ("F)

Figure 4-9 Initial experiment using fluorinated grease in a high vacuum cryogenic friction
experiment.

Solid Lubricant Coating Results

Running an experiment in an application specific geometry under actual loads and sliding

speeds was the initial intent of this project. In these experiments, three coatings were chosen

representing the most temperature sensitive, the least temperature sensitive and an intermediate

sensitivity. The tests were begun at room temperature and the temperature was ramped down to

-600C and back to room temperature throughout the course of the experiment.









0.10-
Test Conditions
.ar O~ 1 Normal Load (Fn)= 100 N
~~O Sliding Speed (v)= 10 mm/sec
(1u Spindle Speed (co)= 20 rpm
S0.05 Vacuum Level = 1x 06 Torr
.9 A MoS2 with Ti
rA gM 8Y Lb 68 d d MoS2 with 5b203 and Au
O MoS2 with Ni
0.0 il*li*
-80 -60 -40 -20 0 20 40
Temperature (oC)

Figure 4-10 Plot of the friction response for all three coatings tested in the experiment. None of
the coatings demonstrated an increase in friction with decreasing temperature in this
geometry. The highlighted outlier in the MoS2 with titanium data set was the last data
point and could indicate the onset of failure for the coating. Many of these coatings
failed after only a few hundred to one-thousand cycles.

The plot shows the friction of all the coatings remained low and constant for the length of

the experiments. The friction coefficient values for all the coatings are consistent with values

found in the cryogenic pin-on-disk experiments. One explanation for the flat friction response

could be high wear rates in the bushing shaft geometry. In the linear reciprocation experiments,

the coatings lasted over 10,000 cycles in all cases. However, in a bushing-shaft configuration the

coating life was often shorter than 2,000 cycles indicating this geometry may be more severe

tribologically. There is some discussion over whether the shaft is riding on an edge of the

bushing or if the system is very axis-symmetric. The wear rate can be estimated using both

assumptions to determine a possible range for the system. In an ideally aligned system where the

shaft contacts the bushing across the entire length of the bushing, the wear rate is estimated at

5x10-6 mm3/Nm. If the system is running on an edge and only a small portion of the shaft is in

contact, the wear rate of the system is 1.55x10-s mm3/Nm. This range of values encompasses

most of the data from the in situ wear experiments; however, it is not possible to determine if

debris generated inside the contact are playing a role in lubricating the contact prior to failure.









CHAPTER 5
DISCUSSION

The term activation energy was introduced by Svante Arrhenius in 1889 to define the

potential energy barrier that had to be overcome before a chemical reaction could take place.

While the potential energy barrier for a reaction does not change as a function of temperature,

the distribution of energy states that each atom or molecule may widen with increasing

temperature. With higher temperatures, it is increasingly likely to find an atom with suitable

energy to overcome the barrier and reach a more stable configuration. A schematic of the

Boltzmann distribution for three temperatures is shown in figure 5-1. This helps to illustrate the

effect temperature has on the system and explains why activated processes occur more readily at

higher temperatures. The example of a chemical reaction is typical when discussing thermal

activation, but there are other phenomena that demonstrate temperature sensitivity. Some

common examples are viscoelastic creep, grain growth and diffusion of an impurity. The last of

these, diffusion of an impurity (interstitial atom) through the lattice structure of another material

is a well documented thermally activated process. It occurs because the atoms are vibrating at

very high frequencies 1013 Hz or higher, and an interstitial atom will make a jump from one site

to another if the orientation of the surrounding atoms is such that it can make the transition based

on its kinetic energy. The amplitude of the vibrations of surrounding atoms, as well as the

kinetic energy of the interstitial atom, are dominated by the temperature of the system. As the

temperature of the system increases, the oscillations of lattice atoms takes them further from the

average position allowing an interstitial atom to squeeze through more easily. This is discussed










in greater detail by Porter and Easterling in Phase Transitions in Metals and Alloys.(53)


Portion of atoms or molecules
with energy states high
enough to overcome the t
potential energy barrier
Activated Energy State

OE

Initial Energy State
T=I 1
T=10 K c



T=100 K m L Final Energy State


Figure 5-1 Illustration of a potential energy barrier typically found in activated chemical
processes. The image to the left of the energy curve is a representation of Boltzmann
energy distribution for a large number of atoms or molecules in a system at different
temperatures. The portion highlighted in black is the area of the curve containing
atoms with energies high enough to overcome the potential energy barrier.

As mentioned before, a previous publication on the friction response of PTFE coatings

initiated these tests. In those experiments, an exponential was used to fit the results. The results

in these experiments seemed to follow a similar trend. The Arrhenius equation was used to

attempt to fit the data. The general form of this equation is shown in (5-1).

-E
k= A-eR-T (5-1)

In this equation, k is the rate coefficient, A is a constant, Ea is the activation energy, R is the

universal gas constant and T is the temperature in Kelvin. The friction response of the coatings

was modeled using this equation in the form shown in (5-2).



pu = u,e R T T (5-2)

Here, Ct is the expected friction coefficient at temperature T, Clo is the room temperature friction

value, the activation energy is represented by Ea, R is the universal gas constant and To is room









temperature in Kelvin. In order to isolate activation energy from the other variable (room

temperature friction coefficient), the friction coefficients are normalized by the room temperature

value.


-= eR (5-3)


A logarithmic transformation is used to create a linear equation in terms of activation energy.

This equation can be fit using a least squares regression analysis.

In (5-4)



c~ I monte-carlo simulation
'' ** .c of activation energy

Simulated field of "
u 1,000 least squares
1000 data points for c
s o regression lines calculated
each mean value *. *
calculated in the 'i-~~l
cyrogenic pin-on- ***
disk experiments. 1, *
.. -Ea=average slope
a u(Ea) = standard deviation
1/T

Figure 5-2 Illustration of methodology for creating a Monte Carlo simulation of activation
energies.


The plot of the raw friction data for MoS2 with Sb203 and the transformed data can be seen in

Eigure 5-2. Because there are only four data points for each curve, it is difficult to determine an

uncertainty in the fit directly. The line represented in figure 5-3(a) and 5-3(b) was determined

using a linear regression technique combined with a Monte Carlo style of simulation. This style

of analysis uses the mean values for each data point (CI and T), shown as circles in figure 5-3, as





























uv l ll ll
-125 -25 75 175
Temperature (*C)


-0.2 -0.1 0 0.1 0.2 0.3 0.4
R1 T1 T1


the central value in the simulation and generates a number of alternative values, in this case


1,000, that could also be reasonable given the standard deviation calculated from the raw data.


Figure 5-3 One data series in standard and logarithmic plot with fit. (a) Example of raw friction
data with fit of the data. (b) Transformed data with plot of fit. This format is used to
generate a fit using least squares regression.


0.25

.~ 0.20

.P 0.15
u


MoS2 &0.25
Ea = 2.6 Id/mol .0
Co= 0.067 I

S0.15


.25

.20

.15


MoS2+Ti 0
a= 5.2 kJ/mol 0
o=0.056 5
8 0


MoS2+Sb20,
Ea= 5.7 kJ/mol
Ro= 0.042


.0

-100 -50 0 50 100 150 200
temperature OC




Test Conditions

Normal Load (Fn) = 2.5 N
Sliding Speed (v) = 20 50 mm/sec
Relative Humidity (())= 0.5%


0 1
-100 -50 0 50 100 150 200
temperature OC
0.25
c ~MoS,+Sb,O,+Gr
.8 0.20 \3 iEa= 8.0kJ/mol
E I o1= 0.034 1
~o0.15 1
.~ 0.10

0.05 3 r


-100 -50 0 50 100 150 200
temperature oC


0
-100 -50 0 50 100 150 200
temperature OC
0.25
: MoS2+Sb20,+Au
-8 0.20 .\Ea = 9.8 kJ/mol
W \.1 po= 0.027



0.05 "z


-100 -50 0 50 100 150 200
temperature OC


Figure 5-4 Activation energy fits for all the MoS2-based composite coatings. This value can be
thought of as a thermal sensitivity for the coating.



For each of the alternative data series, the room temperature friction coefficient and


activation energy are calculated. The mean value of all the fits is the reported activation energy,









and the standard deviation of these values is reported as the uncertainty in activation energy. In

many cases, the spread of the data was sizeable, but the resulting uncertainty in the fit was not.


Table 5-1 Activation energy values from Monte Carlo simulations of all the data along with the
uncertainties in those values.
Coating Material Cco Ea (Id/mol) u(Ea) (k)/mol)

MoS2 with nickel 0.074 1.93 0.45
MoS2 with Sb203 0.042 5.66 0.41
MoS2 with tita nium 0.056 4.09 0.99
MoS2 with Sb203 and graphite (deposited) 0.033 8.18 0.60
MoS2 with Sb203 and gold 0.030 9.50 0.55



The friction results for the MoS2-based coatings in a pin-on-disk configuration show

something about these systems is thermally activated. The increase in friction coefficient with

decreasing temperature varied based on the coating composition. Initially, the hypothesis was

that increasing the amount of MoS2 in the coating was decreasing the temperature sensitivity.

However, after testing several compositions ranging from 50% MoS2 to 95% MoS2, there was

not a strong correlation.

As mentioned before, there was a noticeable difference in the depth of the wear scars and

the appearance of wear debris from one coating to the next. The wear rate values determined in

the in situ wear experiments offered on plausible explanation for the activation energy. By

plotting the of activation energy vs. wear rate for each of the coatings, a correlation between the

two was revealed.









10 g






c 4-





1x10-8 1x10-7 1x10-6 1x10-s 1 x10-4

Wear Rate (mm3/Nm)

SMoS, with Nickel A MoS, with 5b2, ~and Graphite + MoS, with Sb,O,
SMo52 With Titanium MoS2 With Sb20, and Gold

Figure 5-5 Plot of activation energies determined from the cryogenic testing of all the MoS2
coatings as a function of wear rate determined by the in situ experiments for the same
coatings.


The MoS2 with nickel had the highest wear rate of all the MoS2 COmposite coatings. The

wear scar on this film grew beyond the field of view of the SWLI after 7,000 sliding cycles

meaning that wear volumes could not be calculated beyond that point. This sample also had the

lowest sensitivity to temperature (1.93 kJ/mol). The MoS2 with Sb203 and gold wore only one

hundred nanometers over the course of 10,000 cycles leading to the lowest wear rate of all the

coatings tested. This coating also had the highest activation energy of all the coatings (9.5

kJ/mol). As indicated in the plot, there is a trend of decreasing wear rate with increasing

activation energy. The link between these two properties seems to be interfacial sliding. During

interfacial sliding, the interaction between surfaces is dominated by van der Waals forces. In the

MoS2 filmS, the lamellar structure results in planes of the molecule that form sheets with strong

bonds. Between these sheets the sulphur atoms interact creating a weak van der Waals bond










resulting in low shear strength for the material. The commonly used analogy for this coating is

the deck of cards explanation. The layers of MoS2 act like a deck of cards which can be sheared

easily.


Shear Force Shear Force Shear Force

I I
IIl lI



Figure 5-6 The deck of cards illustration for sliding between lamellar MoS2 Sheets. The
interfaces that are shearing change throughout the sliding event, but the shear force
does not change.


In this low shear strength configuration, the only interactions between sheets of MoS2 aef

van der Waals type interactions. It is hypothesized that van der Waals interactions are

temperature sensitive due to a blurring of the potential surface as the temperature of the system

increases. This occurs in a manner similar to that outlined in figure 5-1.

Another interesting result was the increase in wear rate as the MoS2/titanium system was

cycled from a dry nitrogen environment to a humid air environment. This is perhaps due to a

large area of the surface covered with nascent MoS2 that formed during sliding in a dry

environment. This nascent MoS2 likely oxidized when oxygen was reintroduced to the system

causing a large amount of debris generation and an increase in wear volume. The reaction of

MoS2 with 02 is a spontaneous one defined by the following equation.

2M~oS2 + 702 4 2Mo3, + 4SO2 (5-5)

This reaction has a negative change its Gibbs free energy of 2,249 kJ/mol at room temperature

(calculations for all chemical reactions can be found in Appendix D). Initially, the theory was

the water in the system caused the degradation; however, free energy calculations found in

appendix D indicate water will not spontaneously react with MoS2. The untested surfaces are









likely covered with oxides formed after the sample was created, so running in a humid

environment would not cause such a dramatic increase in the wear rate. By sliding in a dry

environment all the oxides formed on the surface are removed initially leaving a low friction low

wear interface to accommodate the remaining cycles. The uncovered MoS2 WOuld be much

more reactive to oxygen causing the surface to oxidize and degrade quickly.









CHAPTER 6
CONCLUSIONS

* Three high-vacuum tribometers were designed and constructed to test samples in a variety
of environments and contact geometries.

* The variable temperature pin-on-disk experiments indicated an increase in friction
coefficient with decreasing temperature.

* There was a strong correlation between activation energy and wear rates for the MoS2-
based coatings.

* The cryogenic bushing tribometer experiment support a hypothesis that if the tribological
contact is dominated by surface deformation and wear, there will be no sensitivity to
temperature variations. The wear rate in this system cannot be directly measured, but
because the system fails after roughly 1,000 cycles, it is a high-wear system.









CHAPTER 7
FUTURE DIRECTIONS

Cryogenic high vacuum linear reciprocating tribometer experiments are in the immediate

future for this proj ect. As mentioned in chapter 2, this tribometer is the easiest to fit with

cryogenic cooling capabilities and resistive heating capabilities; making it an obvious choice for

continuing to evaluate the temperature response of materials. This system has the best chance of

answering the question of why a pin-on-disk experiment in dry nitrogen displayed a strong

temperature dependence and a high vacuum bushing geometry did not. Wear measurements can

be made for this system using the SWLI before the experiment and after the experiment giving a

clear value for volume loss.

The in situ wear experiments proved helpful at determining the behavior of these systems

to environmental changes. There are also many improvements to the technique that can be made

to increase the quality and consistency of the data collected. The first step that can be taken in

future experiments is to have the SWLI take several surface scans of an image and average over

them to improve the repeatability of the scans. Using a single scan, the repeatability of the

measurements was 40nm. The uncertainty of the scans is reduced to 10nm by averaging five

surface scans. This improvement helps eliminate surface noise in the system which is readily

apparent in the line scans used to calculate wear volumes. A second area for improvement is the

calculation of wear volumes. Extrapolating a single line over the entire length of the wear scar

can skew the results, particularly if a piece of debris falls into the area where the line scan is

being taken. Finding a wear volume based on the entire surface scan (usually several hundred

microns in both directions) can give a better representation of the scar. One method for

calculating this value is to use the entire initial surface scan as the undeformed surface and find










the difference between that and a subsequent scan. This should help to eliminate the influence of

noise in the scan and debris on the surface.

Another direction is in the use of molecular dynamics simulations to evaluate the potential

surfaces for model system at different temperatures. This analysis could provide insight as to

why temperature effects are very dramatic in the low temperature regimes and tend to fall off as

the temperature is increased. One theory is the corrugation of the potential surface becomes

blurred as the temperature increases reducing the driving forces necessary to cause shear in the

solid lubricant interface. Careful evaluation of the potential surfaces at varying temperatures

could shed light on this theory.



































K eeto aemesrmnFree Body Diagram
Ff L3=152.4mm r
L2 = 228.6 mm


APPENDIX A
UNCERTAINTY ANALYSIS OF FALEX POD TRIBOMETER

The uncertainty of the friction coefficients measured by the Falex tribometer can be
quantified using the law of propagation of uncertainty. In this case, there are several
contributing factors in the uncertainty of the friction coefficient. They are:
* Mass of the dead weight load (m)

* Ratio of lengths from the dead weight load to the gimbal and the pin contact to the gimbal
(R2) .

* Ratio of lengths from the friction transducer to the gimbal and the pin to the gimbal (R3).

* Friction force measured by the strain gauge (Fs).


L2 = 228.6 mm
(L1 =114.3 mm



m -g rt


L2 = R2L,

3, 3 1,

L,
F, = m-g=
n ~L2


Free Body Diagram
of Dead Weight Loading


m -g

R2


L, R3F,
SL, R,

F, R3F,

F, m -g


x


Figure A-1 Free body diagram of Falex pin-on-disk tribometer with friction coefficient
calculations.

The normal load applied to the sample is calculated based on the ratio R2 and the mass of
the dead weight load.
m-g m-g
F, =
R 2

The law of propagation of uncertainty indicates the uncertainty of this measurement to be:












RZ


o :1: ')(0.05)?


The uncertainty of the friction force can also be calculated using the law of propagation of
uncertainty:

R,

SF SR? IR,
RR F ,:!, F R F
u(F,) = u(F) + R2 ( (


u(F,)= 0.011


The friction coefficient uncertainty value can now be calculated using the uncertainties of its
constituents.






1 -05 F,

u~g) = (0. 1)+ (0.062):
2.5 6.25
u (C)= 0.007

Although the calculated uncertainty of the system is 30% of the measured value at CI = 0.02, the
electronics used to output the data to the data acquisition system report friction coefficient as a
voltage. The resolution of the system is 1 mV corresponding to a friction coefficient of 0.001; a
conservative estimation of the uncertainty of the electronics is 10 mV corresponding to a friction


u(F,)i = "u(mfi + )u(R )


uF,)i = R)u(m) +


u(F,) = ~li5x10-4) +
u(Fn)= 0.062


(o, (0.05) + -(0.05)









coefficient of 0.01. This value will be used as the uncertainty in the measurements for all the
data in the pin on disk experiments.









APPENDIX B
HIGH VACUUM PIN-ON-DISK EXPERIMENT

Environmental Protocols

The high vacuum experiments were meant to evaluate the friction response of the coatings
when run in ambient air (relative humidity ~ 40%) and high vacuum (pressure < lx10-6 Torr).
The first test procedure was to run one track from each sample in ambient air then a second track
at high vacuum. The load used in these tests was 5 N. This value was chosen so coatings with a
low friction coefficient (below 0.05) would still generate a friction force large enough to provide
reasonable certainty in the measurement. These experiments were all run at ambient temperature
(T ~ 230C). The sliding speed for these experiments was approximately 20 mm/sec regardless of
track diameter. This was made possible by the precise spindle control of high vacuum motor.
A second procedure was also used with the high-vacuum pin-on-disk testing referred to as
'pump and purge'. In these tests, the sample was mounted in the pin-on-disk, the chamber was
evacuated to high vacuum and the motor was started. After the sample ran in to steady-state in
high vacuum, the chamber was opened to ambient air at 40% relative humidity. Once the
pressure equilibrated the system was allowed to run to a steady-state condition before the
chamber was evacuated again. These tests were performed to see if cycling environments had
any effect on the friction response of the coatings.

Experimental Procedure

Although the high vacuum pin-on-disk tribometer has multiple sample holders to allow for
a variety of disk diameters, all the samples used in this experiment are 2" diameter '/" thick
disks and '/" diameter balls. The disk sample was placed on the sample mount bolted to the face
of the motor. The cover plate was screwed down over the disk pinching it in place. This design
holds the disk flush to the surface and prevents it from moving during the test. As with the disk
holders, there are multiple pin flexures to accommodate different diameter pin samples. The '/"
sample is inserted into the holder and a set screw is tightened behind the ball holding it firmly in
the housing and preventing it from slipping during the experiment. The sample flexure is
attached to the load cell assembly, and the load cell is tared. This ensures all forces exerted on
the pin sample are accurately read by the transducer. LabView is started and the program written
for this tribometer is executed. The calibration constants for the load cell, pressure gauge and
thermocouples are entered into the software. The location where the data files are to be saved is
chosen and the program waits for the user to begin data acquisition. The micrometer stage the
where the load cell assembly is attached is adjusted to the desired track radius for the
experiment. Once positioned, a second micrometer stage is adjusted to bring the sample into
contact and apply a normal load. The normal force is read in real-time within the LabView
software. When the appropriate load has been applied, the chamber is closed and the system
pumped down to the target vacuum level for the experiment. The data acquisition is started in
the LabView software which subsequently commands the motor to begin. Average values for
ambient pressure, normal force and friction force are recorded throughout the experiment and
appended to an average data file for post processing. Periodically, one complete revolution of
data is recorded.











The sensitivity of MoS2 based coatings to humidity is well documented. Water
contaminates the lamellar structure of MoS2 inCreaSing the stress required to shear the material.
In a solid lubricant coating, this contamination results in an increased friction coefficient. The
expectation for the high vacuum pin-on-disk experiments was that the MoS2 based coatings
would be extremely sensitive to the partial pressure of water in the system. In metallic coatings,
the friction force tends to be dominated by abrasive wear and plastic deformation; there is no
indication that the presence of humidity would improve the ability of these materials to
mechanically deform. The polymeric systems, being chemically inert, were also expected to be
insensitive to humidity. The table below shows all the pin-on-disk results for high vacuum and
humid air.

Table B-1 Summary of results from high vacuum pin-on-disk tribometer in humid laboratory air
and high vacuum. All tests were run at a normal load of 5N and a sliding speed of 20
mm/sec .
Friction Coefficient
Coating Material Air High Vacuum
Hard anodized aluminum 0.75 0.88
Hard anodized aluminum with PTFE 0.13 0.12
Titanium nitride with electroless nickel 0.65 0.61
MoS2 with Sb203 and graphite (burnished) 0.16 0.01
MoS2 with 5b203 and graphite (deposited) 0.13 0.05
MoS2 with titanium 0.12 0.01
MoS2 with nickel 0.27 0.02
PTFE with PEEK and MoS2 0.12 0.1
UHMWPE 0.05 0.03
Near frictionless carbon 0.05 0.008



As expected, the metallic coatings did not display much sensitivity to humidity in the
environment. In all cases, the metallic coatings showed little change compared with the scatter
in the data from these systems. The MoS2 COatings showed dramatic decreases in friction
coefficients in high-vacuum. Again, this result was not surprising. The unexpected results was
in the polymeric systems which displayed decrease in friction response when humidity was
removed from the system. Given the chemical inertness of these systems, it was hard to
determine what caused this change. However, more recent work in this field has indicated there
is a sensitivity to water for PEEK composites.(54)

Pump and Purge Test Results

The other test outlined in the experimental procedures section was the 'pump and purge'
experiment. The importance of this experiment was to determine if repeated cycling of the


Humid Air and High Vacuum Testing










environments had any effect on the friction response of the materials. To evaluate this, two
different samples were chosen, the MoS2 with titanium coating and a bulk polymer composite of
PTFE with PEEK and MoS2. Both of these samples were run in alternating environments to
determine the effects of cycling the environment. The figure below illustrates the results for
each sample.


(a) MoS2 T (b) PTFE + PEEK + MoS2
.... 1000 0.16 1000
0.161 1 100 100io
10 0.141 ). 10* *.~
0.12 -1 Id 1 c 1
-I I 0.1 E '0.12 -0.1
t; 0.01 9 .s 0.01 ,
S0.08 0.10
1E-3 2 1E-3 e
1E-4 0.08 1E-4
0.04-1E-5 -1E-5
1E-6 0.06 A Friion Coefficientl -1E-6
UNL- Ambient Pressure
o.oo 1E-7 1E-7
0 20 40 60 80 100 ( 40 80 120 160 200
Distance (m) Distance (m)

Figure B-1 Pump and purge data results for a thin coating and bulk polymer. (a) MoS2 with
titanium and (b) PTFE with PEEK and MoS2

In both cases, the material began at high vacuum and was exposed to humid air then the
cycle was repeated. The samples displayed a similar response, although the MoS2 with titanium
was much more dramatic, both saw the lowest friction under high vacuum in the initial cycles.
The friction immediately increased upon exposure to humid air, and after pumping the system
back down, the friction did not recover to its initial value, indicating some oxidation or
permanent deformation of the system occurred while running in a humid environment. In the
polymer experiment, it appears the system did not reach an equilibrium state prior to swapping
the environment. The friction was clearly affected by the humidity in the system, but it is not
possible to determine the extent of the change as the friction seems to be increasing continually
during the experiment. One possible explanation for this is a continual deformation of the
surface throughout the test due to the high contact pressures. This results in a more conformal
contact, a larger real area of contact and ultimately higher friction forces.










APPENDIX C
AUGER ELECTRON SPECTROSCOPY OF COATINTGS

Of the coatings tested, two were chosen to have Auger electron spectroscopy performed on
them. These were the MoS2 with titanium coating, mainly because it is designed to be
environmentally insensitive, and the MoS2 with Sb203 and graphite, because this coating is
designed to rely on different constituents to provide solid lubrication in differing environments.
The samples used were from the 10,000 cycle single environment experiments. Two tracks were
used for each sample; the first was run in humid air for 10,000 cycles and the second in dry
nitrogen for 10,000 cycles. In each case, a line scan was performed across the track evaluating
different elements. The MoS2 with titanium samples were checked for carbon, molybdenum,
oxygen, sulphur and titanium. The MoS2 with Sb203 and graphite was tested for carbon,
molybdenum, oxygen, sulphur and antimony. In all four systems there were distinctly different
signatures. The data shown in all the plots are the normalized intensities for each element. This
was chosen because some elements are more likely to emit Auger electrons than others skewing
the raw data. Sulphur, for example, is one of the elements most likely to emit an Auger electron,
so its intensity dominates most of the scans. By normalizing the data, it is possible to look at the
intensities of all the atoms inside and outside the tracks more easily.


0. ,Carbon ,
0 100 200 300 400 500



0.0 Molybdenum
6 160 260 360 460 560




0 100 200 300 400 500
Position (prm)


0. ,Sulphur
0 100 200 300 400 500



0.0
0 100 200 300 400 560
Position (pm)


sldi sped =51 mm/sec ad30p
environment =humid air


Figure C-1 Relative intensities for several elements plotted vs. position across the wear track for
the MoS2 with titanium coating run in humid air environment. Auger data indicates a
high content of titanium and oxygen (likely titania) on the surface inside the wear
track. The signature of the other elements was relatively constant inside and outside
the track.












0.0 a b n , ,


0.0 oydnm
0 100 20 300 00 500

0.0
0 100 200 300 400 500
Poito (m


00,Sulphur
0 100 200 300 400 500



00,Titanium
0 100 200 300 400 500
Position (pm)



sliding speed = 10 mm/sec ad30u
environment =dry nitrogen


Figure C-2 Relative intensities for several elements plotted vs. position across the wear track for
the MoS2 with titanium coating run in dry nitrogen. Like the humid air data, dry
nitrogen data indicates a higher content of titanium and oxygen (likely titania) on the
surface inside the wear track. However, there are also distinct changes in the sulphur
and molybdenum signatures inside the track indicating more MoS2 is on the surface.



The MoS2 with titanium coating had a distinct signature for titanium and oxygen inside the
wear track in both humid air and dry nitrogen. This indicates titania is being formed in the
system as the track is worn in. The signals from molybdenum and sulphur remained constant
across the track indicating there was still a reasonable amount of MoS2 remaining in the track.
The track run in dry nitrogen did have an increase in titanium and oxygen inside the track, but
there was also an increase in sulphur. While the environment is dry nitrogen, there are always
impurities present. These results indicate any oxygen in the environment seems to be trapped by
the titanium in the coating, but the presence of titania does not impede the MoS2 and may help to
prevent its oxidation.


1.0
0.5

0.0 Oxgn
0 100 200 300 400 500
Poito (m


0 100 200 300 400 500


Position (pm)
Test Conditions Wear track located
normal load = 5 N between 150 plm
sliding speed = 10 mm/sec and 350 pm
environment =humid air


Figure C-3 Relative intensities for several elements plotted vs. position across the wear track for
the AFRL chameleon coating run in humid air. There is no signal from the
molybdenum in this coating. Oxygen and carbon are primarily found inside the wear
track while sulphur and antimony are depleted inside the track.





Test Conditions
normal load= 5 N
sliding speed = 10 mm/sec
environment =dry nitrogen


Wear track located
between 150 pm
and 350 pm


Figure C-4 Relative intensities for several elements plotted vs. position across the wear track for
the AFRL chameleon coating run in dry nitrogen. The molybdenum, sulphur and
carbon intensities increase inside the wear track. The difference between this scan
and the open air tests indicate MoS2 is, in fact, drawn to the surface of the coating
during wear.



The MoS2 with Sb203 and graphite results indicated that the coating does have a stronger
signal from MoS2 On the surface running in dry nitrogen as opposed to humid air. The humid air
results indicated mostly carbon and oxygen in the wear track. This result is expected based on
the hypothesis that the preferred solid lubricant is drawn to the surface. The lack of a signal
from molybdenum and the sharp decrease in sulphur signal inside the wear track also confirm
that carbon (could be graphitic), not MoS2, iS the primary constituent on the surface of the
material after running in a humid environment.


1.0 1.0

0.505

0.0 Molybdenum 0.50 Aimo
0 100 200 300 400 500 0 100 200 300 400 500



10 Position (pm)


0.5 Oxygen
0.0 ,
0 100 200 300 400 500
Position (plm)









APPENDIX D
GIBBS FREE ENERGY CALCULATIONS

The oxidation of MoS2 to MoO3 is a process of importance in this study. Determining
which species are primarily involved in the reaction is key to predicting when it will occur. To
that end, the two species present in humid air that are known to react with a variety of materials,
Oxygen and water were evaluated as causes for oxidation.

A chemical reaction involving water and MoS2 is listed below. This equation is generated
solely by balancing the reactants with the products.


M~oS2 + 3H20 4 MoO3, + 2H2S + H2


(D-1)


For this chemical reaction to occur spontaneously the Gibbs free energy of the system must
be less than 0 kJ/mol. The reaction is assumed to occur at room temperature 298K. The values
for enthalpy and entropy for each species is listed in the table below.

Table D-1. List of values for enthalpy and entropy for species in (D-1).(55)
Species Enthalpy Entropy


MoS2
H20
MoO3
H2S
H2
02
SO2
H2SO4


-276 k)/mol
-242 k)/mol
-745 kJ/mol
-21 kJ/mol
0 k)/mol
0 k)/mol
-297 kl/mol
-735 kJ/mol


63 J/mol-K
189 J/mol-K
78 J/mol-K
206 J/mol-K
131 J/mol-K
205 J/mol-K
248 J/mol-K
299 J/mol-K


To calculate the free energy of the equation, the value can be found using equation D-2.

AH + AS T = G

The value for the reaction listed in D-1 is calculated as follows:


Huoo34+ 2HH2S +HH2)-(H Hs+ 3HH20

1000 [Sioo + 2SH2S+ SH2 Sno + 3SH21= G


(D-2)


(D-3)









(-745-42 -0)-(-276 -726) +
298 (D-4)
[(78+412+131)-(63+563) = G
1000


kIn
G = 205 (D-5)
mol

This result indicates the reaction will not occur spontaneously until the temperature is far above
the decomposition temperature of the material.

Another possible reaction is that Oxygen present in the environment is causing the
oxidation of the MoS2. This reaction is defined in D-6.

2M~oS2 + 702 4 2Mo3, +4SO2 (D-6)

Using D-2, and the values from Table D-1, the free energy of this reaction is calculated as
follows:
(2Huoo3 +4Hso2) -2Huos2 +7Ho2
T (D-7)
1000 [2Snoo3 +4~So2)- 2Snos2 +7So21= G

(-1490 -1188)-(-552 -0)+
T (D-8)
10 (156+992)-(126+1435) = G

kIC
G = -2249 (D-9)
mol

Because the free energy of the reaction is far below zero at room temperature, the reaction is
expected to occur spontaneously.

A third reaction includes the combination of water and oxygen in the system to oxidize the
MoS2.

2M~oS2 + 902 + 4H2 4 2Mo3, + 4H2SO2 (D-10)

Using the same method from D-7 to D-9, the free energy of this reaction is calculated to be

kJ
G = -4177 (D-11)
mol









Both water and oxygen can be found in the system in all the experiments listed in this study.
These results only strengthen the conclusion that further testing must be performed in vacuum
where the partial pressures of every species can be accurately measured.










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BIOGRAPHICAL SKETCH

Matthew Adam Hamilton was born January 21, 1978 in Orlando, Florida. He spent a

maj ority of his youth in Bradenton, Florida where he attended Southeast High School. In August

of 1996, he entered the University of Florida. Subsequently, the football team won its first

national championship. A turning point in his academic career occurred in the year 2000, when

he began doing research with Dr. W. Gregory Sawyer in the Tribology Laboratory. After

earning his bachelor's degree in May 2001, Matthew continued with graduate studies in the field

of tribology. Upon completing his master' s degree in May 2003, he took a j ob with a software

company in New York City. After one year, the opportunity to come back to school and

complete his doctorate was presented to him by Greg Sawyer. In June 2004, Matthew was

readmitted to the University of Florida. He completed his doctorate in June 2007 accompanied

by another national championship for the football team and back-to-back national championships

for the basketball team. With aspirations of becoming a professor, Matthew accepted a post-

doctoral appointment with Dr. Robert Carpick at the University of Pennsylvania to begin in

August 2007.





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EVALUATION OF TRIBOLOGICAL RESPONSE OF MOLYDENUM DISULPHIDEBASED COATINGS TO VARYING ENVIRONMENTS By MATTHEW A. HAMILTON A DISSERTATION PRESENTED TO THE GRADUATE SCHOOL OF THE UNIVERSITY OF FLOR IDA IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF DOCTOR OF PHILOSOPHY UNIVERSITY OF FLORIDA 2007

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2 2007 Matthew A. Hamilton

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3 For Hobbes

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4 ACKNOWLEDGMENTS I would like to thank Dr. W. Gregory Sawyer for his help and guidance throughout this undertaking; without him, this di ssertation would never have been realized. I would also like to thank my friends and family for their input, enco uragement and comic relief. Finally, I thank my wife Melissa. Her support, above all others, has allowed me to achieve this goal in my life.

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5 TABLE OF CONTENTS page ACKNOWLEDGMENTS...............................................................................................................4 LIST OF TABLES................................................................................................................. ..........7 LIST OF FIGURES.........................................................................................................................8 ABSTRACT...................................................................................................................................11 CHAP TER 1 INTRODUCTION................................................................................................................... ..13 1.1 Motivation of Research..................................................................................................... 13 1.2 Literature Review.......................................................................................................... ...14 2 EQUIPMENT DESIGN............................................................................................................. 20 2.1 Design Philosophy............................................................................................................20 2.2 High Vacuum Chamber.................................................................................................... 21 2.3 High Vacuum Pin-On-Disk Tribometer........................................................................... 23 2.4 High Vacuum Linear Reciprocator...................................................................................27 2.5 Cryogenic Pin-On-Disk Tribometer................................................................................. 30 2.6 In Situ Wear Tribometer................................................................................................... 33 2.7 Cryogenic High Vacuum Bushing Tribometer................................................................. 36 3 VARIABLE ENVIRONMEN T EXPERIMENTS.................................................................... 44 3.1 Overview of Materials Used in Experiments.................................................................... 44 3.2 Cryogenic Pin-On-Disk Experiment................................................................................. 45 Environmental Protocols................................................................................................. 45 Experimental Procedure.................................................................................................. 47 3.3 In Situ Wear Experim ent.................................................................................................. 49 Sample Description.........................................................................................................50 Environmental Protocols................................................................................................. 51 Experimental Procedure.................................................................................................. 52 3.4 Cryogenic High Vacuum Bushing Experiment................................................................ 55 Environmental Protocols................................................................................................. 55 Material Preparation........................................................................................................56 Experimental Procedure.................................................................................................. 56 4 RESULTS........................................................................................................................ ..........58 4.1 Cryogenic Pin-On-Disk Experiment................................................................................. 58 Determining a Friction Coefficient................................................................................. 58

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6 Experimental Results.......................................................................................................58 4.2 In Situ Wear Experim ent.................................................................................................. 64 4.3 Cryogenic High Vacuum Bushing Experiment................................................................ 68 Preliminary Experiment................................................................................................... 68 Solid Lubricant Coating Results...................................................................................... 69 5 DISCUSSION............................................................................................................................71 6 CONCLUSIONS........................................................................................................................79 7 FUTURE DIRECTIONS...........................................................................................................80 APPENDIX A UNCERTAINTY ANALYSIS OF FALEX POD TRIBOMETER.......................................... 82 B HIGH VACUUM PIN-ONDISK EXPERIMENT...................................................................85 Environmental Protocols................................................................................................. 85 Experimental Procedure.................................................................................................. 85 Humid Air and High Vacuum Testing............................................................................ 86 Pump and Purge Test Results.......................................................................................... 86 C AUGER ELECTRON SPECTROSCOPY OF COATINGS.................................................... 88 D GIBBS FREE ENERGY CALCULATIONS........................................................................... 91 LIST OF REFERENCES...............................................................................................................94 BIOGRAPHICAL SKETCH.........................................................................................................98

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7 LIST OF TABLES Table page 3-1 List of coatings initially te sted for environm ental sensitivity................................................. 44 4-1 All the steady state wear testi ng in hum id air and dry nitrogen.............................................. 68 5-1 Activation energy values fr om Monte Carlo simulations of all the data along with the uncertainties in those values..............................................................................................75 B-1 Summary of results from high vacuum pin-on-disk tribom eter in humid laboratory air and high vacuum. All tests were run at a normal load of 5N and a sliding speed of 20 mm/sec...................................................................................................................... ....86 D-1 List of values for enthalpy and entropy for species in (D-1).( 55 )........................................91

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8 LIST OF FIGURES Figure page 2-1 Vacuum chamber pump down curve with inset of residual gas analysis results at high vacuum ...............................................................................................................................23 2-2 Pin-on-disk design variations.............................................................................................. ....25 2-3 Components of the high vacuum linear reciprocating tribom eter........................................... 29 2-4 Schematic of modified pin-on-disk tribom ete r used in the temp erature studies. The forces in the system are depicted in the image above using grey arrows.......................... 31 2-5 Cryogenic pin-on-disk sample housing...................................................................................32 2-6 Overview of in situ tribometer............................................................................................. ....34 2-7 Bushing tribometer........................................................................................................ .........37 2-8 Illustration of load path to ground for bushing tribom eter design.......................................... 39 2-9 Friction coefficient derivation for bushing tribom eter results................................................. 42 3-1 Description of impinging jet technique used to cool and clean th e surface of the disk during experiments............................................................................................................. 47 3-2 Screenshot of LabView software writ ten for the in situ wear experim ents............................. 53 4-1 Friction plot of the electro less nickel coating at -25C. The plot indicates the large fluctuations in friction coe fficient throughout the experim ent. The average value for this data was = 0.8 with a standard deviat ion of 0.1. The temperature of the system is also plotted to indicate the small fl uctuations in temperature over a 15 minute period.................................................................................................................................59 4-2 Friction response of metallic coatings to v arying temperature. The friction forces were scattered due to the large amount of su rface deformation and wear that occurred during the test.....................................................................................................................60 4-3 Friction response of metallic coatings with solid lubrican t impregnated. (a) The friction results for hard anodize with PTFE coating at -30C. (b) The friction results for the metallic coatings impregnated with PTFE were much more uniform than the metallic coatings alone, but the coatings did not appear to be sensitive to changing temperatures................................................................................................................... ....61 4-4 Results for bulk polymeric samples against a stain less steel pin. The PTFE-based samples displayed a linear decrease in friction coefficient with increasing

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9 temperature, while the UHMWPE sample maintained a constant friction coefficient regardless of temperature................................................................................................... 62 4-5 Friction coefficients of MoS2 based solid lubricant coatings at varying temperatures. The nearly pure MoS2 with nickel coating (95% MoS2, 5% Ni) had the lowest sensitivity to temperature. The commercially available MoS2 with Sb2O3 and gold had the lowest recorded friction at 180 C and the highest increase in friction as temperature decreased (over an order of magnitude)........................................................ 63 4-6 Methodology for calculating wear volume. (a) Plot of a line scan taken from the undeformed surface scan made by the Zygo prior to beginning the experiment and a line scan taken from the surface scan made after 6,000 cycles. (b) Estimation of the wear scar cross-sectional ar ea using the difference between the two line scans. This area is multiplied by the length of the track to estimate a wear volume............................ 65 4-7 Plot of volume lost vs. work input into the system The plot indicates the wear of this coating is sensitive to the partial pressure of water in the environment. In this case, the wear rate initially in air is much less severe than transitioning from a dry environment to a humid one.............................................................................................. 66 4-8 Plot of volume lost vs. work for the MoS2 with Sb2O3 and gold coating. The uncertainty in this value is nearly 100% of the valu e due to the small volume loss over the course of the entire 2,000 cycles. This result led to testing at hi gher normal loads and larger numbers of cycles.............................................................................................................. 67 4-9 Initial experiment using fluorinated grease in a high vacuum cryogenic friction experiment..................................................................................................................... .....69 4-10 Plot of the friction response for all three coatings tested in the experim ent. None of the coatings demonstrated an in crease in friction with decr easing temperature in this geometry. The highlighted outlier in the MoS2 with titanium data set was the last data point and could indicate the onset of failure for the coating. Many of these coatings failed after only a few hundred to one-thousand cycles......................................70 5-1 Illustration of a potential en ergy barrier typically found in ac tivated chem ical processes. The image to the left of the energy curv e is a representation of Boltzmann energy distribution for a large number of atoms or molecules in a system at different temperatures. The portion highlighted in black is the area of the curve containing atoms with energies high enough to ov ercome the potential energy barrier..................... 72 5-2 Illustration of methodology fo r creating a Monte Carlo simula tion of activation energies. ... 73 5-3 One data series in standard a nd logarithm ic plot with fit. ..................................................... 74 5-4 Activation energy fits for all the MoS2-based composite coatings. This value can be thought of as a thermal sensitivity for the coating............................................................. 74

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10 5-5 Plot of activation energies determined f rom the cryogenic testing of all the MoS2 coatings as a function of wear rate dete rmined by the in situ experiments for the same coatings.................................................................................................................. ...76 5-6 The deck of cards illustration for sliding between lamellar MoS2 sheets. The interfaces that are shearing change th roughout the sliding event, but the shear force does not change................................................................................................................................77 A-1 Free body diagram of Falex pin-on-disk tr ibometer with friction coefficient calculations........................................................................................................................82 B-1 Pump and purge data results fo r a thin coating and bulk polym er.......................................... 87 C-1 Relative intensities for several elements pl otted vs. position across the wear track for the MoS2 with titanium coating run in humid air environment. Auger data indicates a high content of titanium and oxygen (likely titania) on the surface inside the wear track. The signature of the other elements was relatively constant inside and outside the track..............................................................................................................................88 C-2 Relative intensities for several elements pl otted vs. position across the wear track for the MoS2 with titanium coating run in dry nitrogen. Like the humid air data, dry nitrogen data indicates a higher content of titanium and oxygen (l ikely titania) on the surface inside the wear track. However, ther e are also distinct changes in the sulphur and molybdenum signatures inside the track indicating more MoS2 is on the surface...... 89 C-3 Relative intensities for several elements pl otted vs. position across the wear track for the AFRL cha meleon coating run in humid air. There is no signal from the molybdenum in this coating. Oxygen and carbon are prim arily found inside the wear track while sulphur and antimony are depl eted inside the track...........................................................89 C-4 Relative intensities for several elements pl otted vs. position across the wear track for the AFRL cha meleon coating run in dry nitrog en. The molybdenum, sulphur and carbon intensities increase inside the wear track. The difference between this scan and the open air tests indicate MoS2 is, in fact, drawn to the surface of the coating during wear....................................................................................................................................90

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11 Abstract of Dissertation Pres ented to the Graduate School of the University of Florida in Partial Fulfillment of the Requirements for the Degree of Doctor of Philosophy EVALUATION OF TRIBOLOGICAL RESPONSE OF MOLYDENUM DISULPHIDEBASED COATINGS TO VARYING ENVIRONMENTS By Matthew A. Hamilton August 2007 Chair: W. Gregory Sawyer Major: Mechanical Engineering Molybdenum disulphide-based coati ngs are the standard solid lubr icant coatings for mechanisms intended to operate in low earth orbit. These co atings are known to provide low friction in a high-vacuum environment. In my study, a variety of hard metallic and so lid lubricant coatings were tested in a self-mated configuration at temperatures ranging from -80C to 180C to evaluate the friction response to varying temperature. The test s were performed using a pin-ondisk tribometer inside an environment chamber at fixed normal load and sliding velocity. Systems where the contact was dominated by plas tic deformation and wear debris generation showed no evidence of thermal sensitivity and ha d friction coefficients ranging from 0.2 to 0.9. However systems where motion was accommodated by interfacial sliding experienced as high as an order of magnitude increase in friction coefficient (0.02 to 0.2) as the temperature of the system was decreased from 180C to -80C. The low friction coefficients of many solid lubrican t coatings have been attributed to interfacial sliding in the material. In these experiments, th e goal was to prove the systems were interfacially sliding by analyzing the evolution of surface to pography. The results indicated that for the MoS2-based coatings it took an average of 30 cycles to remove a single atomic layer of MoS2 (6 angstroms). This showed that interfacial slidi ng was occurring in a majo rity of the contact.

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12 The end goal of the study was to comment on the application specific performance of the coatings in a bushing-shaft configuration. In this system, the fricti on coefficient showed athermal behavior caused by severe wear in the contact area. Ultimately designers of space mechanisms will have to determine the severity of the tribological contact in order to estimate the fluctuations in friction coefficient that may be experienced by varying temperature.

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13 CHAPTER 1 INTRODUCTION 1.1 Motivation of Research Designing mechanisms to perform in low ear th orbit requires predictable performance from the sliding interfaces. There are several challenges facing designers of space mechanisms. First, the temperatures that may be experienced in orbit vary widely (-200C to 200C). Second, the pressures found in low earth orbit are belo w the limit for many traditional lubricants (1x10-9 Torr). Finally, the mechanisms are typically tested on earth prior to launch, so materials chosen must not only function in both e nvironments but also resist oxidation and degradation due to ambient species. Any one of these variables ca n have an effect on th e tribology of a sliding interface. Characterizing these effects and providing models for predicting performance under these conditions is the focus of this research. Many designers characterize frictional losses in a system as a function of the mechanics of the system. While the mechanics can drive some of these losses, the environment can play an equally important role in the re sponse of the system. The influence of temperature on the friction of a system can be dramatic. Polytetraf luoroethylene (PTFE) is a primary example of a commonly used solid lubricant that is affected by temperature changes. In this polymer, temperature changes can influence both its mechanical and tri bological properties.( 1-4) Molybdenum disulphide (MoS2) has become a standard solid lubricant for high vacuum or dry conditions. The interest in the tribology of MoS2 can be seen in the large number of publications focused on this topic. There are howe ver, two areas that have not been exhaustively studied. The first is the macr oscopic friction response of MoS2 coatings to decreasing temperatures. The second is an in situ quantification of wear for MoS2 and MoS2-based composite coatings while cycling environment. A review of current lite rature yields only one

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14 publication using MoS2 in a cryogenic environm ent, but this paper was simply a publication of test results and offered no hypothesis for the results found.( 5) A paper by McCook et al. looked at PTFE coatings at reduced temperatures and posed the hypothesis that friction of self-mated PTFE may be thermally activated. The group also notes the appearance of a cutoff temperature below which the friction no longer rises, but levels out.(6) A more recent work on highly oriented pyrolytic graphite (HOPG) also noted this type of th ermally activated friction response including a temperature limit, however, this wo rk was on a much smaller scale in ultra-high vacuum.( 7) Another study performed by Schirmeisen et al. found this type of increasing friction for untreated silicon (111) s liding against a silicon tip.( 7, 8 ) Like the experiment by Zhao et al. this experiment was performed at ultra-high v acuum (UHV) using an atomic force microscope (AFM). The results indicate there may be a physical limit where interf acial sliding of these materials is no longer the most convenient path for energy dissipation. However, it may be just as possible that the cutoff in each case is caused by separate phenomena. Therefore, the goal of this study was to determine if in terfacial sliding of self-mated materials displayed a thermally activated behavior and to identify the transition from this behavior to athermal response at the macroscale. 1.2 Literature Review Brief History of Molybdenum Disulphide Coatings As early as 1941 there is reference to the low friction of molybdenite in vacuum. It was discovered in a search for a replacement lubricant for the rotating anode x-ray tube.( 9) One of the first substantial studies of MoS2 as a solid lubricant was pe rformed by Johnson, Godfrey and Bisson in 1948. They looked at th e effects of contact pressure and sliding speed on the friction coefficient for a thin MoS2 coating on steel in air. One conclusion drawn from the study was MoS2 is very effective in reducing fric tion at high sliding velocities. This film material was very

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15 tenacious, was chemically and thermally stable a nd consequently should have many practical applications .( 10) After the evaluation of the steady-st ate tribological prope rties in air, the sensitivity of friction coefficient of MoS2 to buildup of surface oxide layers and crystallite orientation was reported by Johnson a nd Vaughn in 1956. This study held MoS2 pins stationary for varying periods of time and observed the star ting friction coefficient and steady-state friction coefficient. They determined that the bu ildup effect caused an in crease in the friction coefficient for a short time (less than 10 minutes) befo re it returned to a steady-state value. Once the system had run to steady state in one directio n, the motor spindle was reversed to analyze the effect of crystalline orientation on friction. It was determined that orientation had no effect on steady-state friction.( 11) The popularity of the new material led to expansion of its testing. Deacon and Goodman examined the role of el evated temperature on the friction of MoS2 and several other lamellar solids. A pin-on-disk style of tribometer was used to measure the friction of a diameter hemisphere of platinum sliding on a flat disk of platinum both surfaces covered with a thin MoS2 layer. The flat disk was heated from room temperature up to 600C and friction coefficient was initially above 0.2 but quickly ran down to a value of 0.07 for temperatures from 40C to 300C at which point it steadily rose to a value of 0.5 at 530C. The results indicated MoS2 could maintain a friction coefficient less than 0.1 up to 300C where the molecule began to oxidize to MoO3 in air.( 12) Because of the distinct difference in the friction of MoS2 in air compared with high-vacuum Haltner a nd Oliver looked at the effects of the partial pressure of water on the friction response of MoS2. Pellets of MoS2 were run against a copper substrate to create a transfer film in air and friction respons e was observed while varying the humidity surrounding the sample from ambient leve ls 25% RH to 0.5% RH. This study posed a model for friction coefficients as a func tion of increasing partial pressure of H20 in the

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16 system.( 13) Further environmental studies were done with MoS2 in ultra-high vacuum (p < 10-9 Torr) where the temperature was elevated to over 930C before thermal decomposition was reached.(14) Ross and Sussman found the oxidation of MoS2 to MoO3 to be a self-limiting process; meaning the formation of a monolayer of oxide prevented further oxidation. The study showed steep initial increase in the percentage of MoO3 when exposed to humid air, but as the exposure time increased, there was no increase in MoO3.( 15) After the discovery of the low friction and thermal stability of MoS2 films it became a primary candidate for many aerospace applications particularly elevated temperature systems such as jet engines. A study of the friction of the available MoS2 coatings at temperatures ranging from -75C to 200C was published by H opkins and Campbell in 1969. A majority of the coatings in this experiment were MoS2 with graphite using diffe rent binders (epoxy, sodium silicate, etc.). The focus of the study was to de termine how different binder materials affected the friction and wear performance of MoS2 films. The coatings were applied using an air brush then cured onto the surface. The data from these tests indicated a tre nd of decreasing kinetic friction with increasing temperature.( 5) Other studies focused on the wear life of burnished coatings in an effort to improve the life expectancy of the coating.( 16) Up to this point, a majority of the studies used manual deposition such as burnishing, pain ting or evaporating the material onto the sample. A number of groups were also working on improving deposition techniques for MoS2. The performance of sputtered MoS2 films was first studied by Talivaldis Spalvins. His studies showed an increased wear life and a more stable fr iction response over the life of the sample for a sputtered MoS2 films in vacuum on a variety of substrates.( 17, 18 ) One important conclusion that appeared in most of the papers on deposited films was cleanliness of the substrate heavily influenced the performan ce of the film. Another discovery, documented by

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17 Petrov et al. was the effect of the substrate temperature deposition on the degree of crystallization and the evolution of the structure inside th e film during growth.( 19) The study by Holinski and Gansheimer on the lubricating mechanism of MoS2 indicated the layered structure shears easily and covers rough asperities on th e surface. The inter actions between MoS2 layers are weak allowing for low shear strength and low fr iction coefficient. This study illustrates the sliding as a deck of cards where multiple layers participate in the shearing events simultaneously. Using SEM, the authors were able to quantify th at roughly 25 atomic layers participated in sliding during the bearings tests performed.( 20) Finally, the crysta llinity of the MoS2 coating was interrogated. This study i ndicated that an amorphous MoS2 had poor tribological properties, in particular a high friction coefficient ( = 0.4) and acted as an ab rasive that wore the counterface. Spalvins also studied the wear process for MoS2 films that grew in finger-like structures from the substrate. Th e results indicated that these stru ctures break early in the wear process leaving a well-adhe red thin layer of MoS2 bonded to the surface to accommodate the shear.( 21 ) The success of new deposition techniques (e.g. radio frequency (RF) magnetron sputtering( 22 ), DC-diode sputtering( 23) and RF diode sputtering( 24 )) gave researchers the ability to form composite coatings and study the effects of other constitu ents added to the MoS2 films. Initially the hope was to reinforce the MoS2 and extend its life, but as new materials were introduced and deposition techniques were improved another goa l arose; to combine other materials with MoS2 that could help prev ent oxidation of MoS2 and provide solid lubrication for environments where MoS2 was not the best choice (namely humid air). This goal has remained a focus for many groups developing high performance so lid lubricant films over the past 30 years. Zabinski et al. studied the morphological and tribological properties of MoS2 films codeposited

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18 with a variety of materials to identify changes that accompan ied improved friction and wear characteristics. In this study, gold, iron, nickel and antimony trioxide were used as dopants for MoS2 films. They concluded a decrease in crystalline grain size decreased the friction coefficient and improved wear resistance of the coatings. In the study a composite of MoS2/Sb2O3/Au resulted in the smallest grain sizes of all the coatings.( 25 ) Further studies under increased temperatures led to the development of a yttria-stabilized zirconia (YSZ), gold, MoS2, diamond-like carbon (DLC) composite termed a chameleon coating. This coating was developed at the Air Force Research Labs (AFR L) and produced a friction coefficient of 0.10 in air reportedly due to the graphitic carbon from DLC and 0.02 in a dry nitrogen environment because of MoS2. As the temperature of the system was elevated to 500C, sliding was accommodated by drawing gold to the surface yielding a friction coefficient of 0.15.( 26) Many other studies have been performed using MoS2 in combination with other constituents in an attempt to further improve performance over an increasing range of operating temperatures and environments.( 27-31) Other developments in coating technologies are the use of com posites to control the nanocrystalline structure of the coatings. Th e AFRL has performed a number of studies on the structure of coatings using diffe rent recipes and correlated that with tribological performance of the films.( 26, 32-34) Another improvement in thin film coatings is customizing the surface prior to applying the coatings. By designing the surf ace topography, it is possi ble to control how the surface will wear and ensure that the contact is never starved of lubricant. One technique that has gained popularity is texturing the surface prio r to applying coatings. Using a laser, micronsized dimples are created on the surface of a solid created. These holes act as reservoirs for the solid lubricant so it is replenished as the surf ace wears. This technique, developed by Andre

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19 Voevodin at AFRL, has had success at maintain ing low friction for extended numbers of cycles.(35) Although the use of deposition techniques has added diversity to the materials chosen to be added with MoS2 for new coatings, it has also been show n to negatively influence the quality of the MoS2. Several studies have been performed on testing the quality of MoS2 in deposited coatings and found that the background pressure of the deposition chamber drastically affects the ratio of Sulphur to Molybdenum in the coatings For deposited coati ngs where the chamber pressure is on the order of 1x10-4 to 1x10-3 Torr, the ratio for S to Mo in the resulting coatings ranges from 1.5 to 1.9. These coatings demonstrated a range of friction co efficient values from 0.007 to 0.1 in inert environments.( 36) However, in systems where deposition occurs under ultra-high vacuum (1x10-9 Torr or less) the ratio is 1.97.( 37) The very pure MoS2 coatings had recorded friction values below 0.002 in vacuum.( 36) The poor quality of the sputtered MoS2 leads to Oxygen substitution in th e lattice structure and edge sites throughout the coating that are primary nucleation points for oxidation to MoO3.( 38)

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20 CHAPTER 2 EQUIPMENT DESIGN 2.1 Design Philosophy The goal in the design of the high vacuum tribometers was to create devices that functioned inside a vacuum chamber, but required no chamber in particular. In other words, the tribometers could not require m echanical feedthroughs. There were two reasons driving this decision initially. The first was that only one vacuum chamber was being fabricated, and to keep that chamber as simple as possible, no mechan ical feedthroughs could be required. The second was that keeping the motor for the system inside the chamber as near to the sample as possible minimizing immeasurable losses that arise when complex loading and motion systems are used to accommodate these feedthroughs. Driving th e systems with vacuum compatible motors simplified the designs guaranteeing that all loads reacted by the sample are read by the load cell in each system. This design decision requires a moto r and stage system that can reliably oscillate for millions of cycles without adjustment. All the other components can be adjusted prior to evacuating the chamber; however, a moto r failure ruins the experiment. The philosophy of connecting a lo ad cell directly in the pa th between the sample and ground was put forth by Schmitz et al. as one th at minimizes the uncer tainties in the force measurements.( 39, 40) This philosophy is followed in all the custom-built tribometers discussed here. To maximize the sensitivity of this tribomet er, the load cell is mounted so the friction force is measured by the most sensitive axes of the load cell. In some cases, the load cell is oriented so the normal force is also measured using the more sensitive axes, particularly in cases where the designed normal load is small compared to the capacity of the load cell. During the construction of the tribometers, th e software written for each one was equipped with the ability to record phase-locked data. This means the positional data is recorded

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21 simultaneously with normal load and friction force. This data is useful for debugging issues with the design and can expose structure in the data that is caused by sync hronous topography in the sample. 2.2 High Vacuum Chamber The ambient pressures at 20 km and 2,000 km are 10 Torr and 1x10-15 Torr, respectively. In a vacuum system, pressures above 1x10-3 Torr can be maintained th rough a controlled leaking of gas into the system. Ultra-high vacuum levels (< 10-9 Torr) are typically obtained through metal sealed stainless steel chambers that are baked to remove water and continuously pumped. A decision to operate at high vacuum levels (< 10-6 Torr) was made in effort to minimize complexity and maximize sample throughput. Because of gas su rface interactions ( 41 ) that can often cause dramatic changes in the frictional be havior, the vacuum level or level of cover gas cleanliness required to simulate space environments continues to be discussed and is likely a function of sample material, geom etry, and sliding speed (some authors suggest that pressures as low as 10-3 Torr may be sufficient for certain materials ( 32, 42 )). Controlling the environment around the tribometer is the most critical aspect of vacuum tribology equipment. For this tribometer a cu stom-built stainless steel vacuum chamber was designed to reach the base pressure of 1x10-6 Torr in less than two hour s. The transition from viscous flow to molecular flow (occurring ~ 1x10-3 Torr) requires a pump for each regime. A dry-scroll pump was selected for the viscous fl ow regime (low vacuum) because of its high pumping speed and it oil free operation, which ensure s there can be no back streaming of oil that would contaminate the chamber.( 43) A cryogenic vacuum pump is used after the crossover to the molecular flow regime (high-vacuum). This system works using a liquid helium compressor to cool a series of surfaces inside the pump from 77 K to 4 K. The fundamental theory behind cryogenic vacuum pumps is that desorption of chemisorbed and physisorbed molecules on a

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22 surface is exponentially dependent on temperature( 44). There are a number of reasons that this type of high vacuum pump was selected. First with a turbomolecular pump the system can be destroyed by a sudden increase in pressure. Thes e pressure rises could be caused by a leak or a valve being opened mistakenly. The increase in pressure on one side of the system of turbine blades causes them to deflect enough that they will collide with the second level of blades resulting in catastrophic failure. However, if this occurs in a cryogenic pump, the pump will be saturated and can no longer pump gases. The solution to this is to empty or regenerate the pump by allowing it to warm to room temperature agai n. The second reason is the pump system works using cryogenic stages so water is pumped at a much higher rate than with a turbomolecular pump. Although the pumping speeds for air and nitr ogen are equivalent be tween similarly sized cryogenic and turbomolecular pumps (1500 L/sec), the pumping speed for water is 266% higher for the cryogenic pump (4000 L/sec) than it is for the turbo pump (1500 L/sec) In stainless steel vacuum chambers, water is often the primary s ource of background pressure for the system. The pump-down curve for the vacuum system is pl otted in Figure 2-1, along with the species found in the chamber as reported by the residual gas analyzer. Clearly th e partial pressure of water is still the largest contributor of all the species found at this pressure. In an effort to drive as much adsorbed water as possible from the interior su rfaces, heater jackets were wrapped around the external surface of the vacuum ch amber; this helps to helps to drive water off the surface during pump-down and helps to reduce th e rate of water adsorption when the chamber is vented. The background pressure is read by a pr essure gauge mounted directly to the chamber, which outputs a 0V to 10V signal indicating th e pressure reading. The system can also determine ambient species present inside the chamber by using a residual gas analyzer (RGA). The RGA works by ionizing a sample of the molecules present in the chamber and measuring their mass to charge

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23 ratio. This ratio is often unique to a single molecule or group of molecules and offers insight into the molecules present due to the environment and those coming from the sample itself. Figure 2-1 Vacuum chamber pump down curve with inset of residual gas analysis results at high vacuum. 2.3 High Vacuum Pin-On-Disk Tribometer Mimicking a space environment inside the laboratory presents multiple challenges, one of the most debated issues is the vacuum level requ ired to accurately reproduce space performance. There is no industry standard for this value, but the Harris Corporation in dicated that a vacuum level of 10-6 Torr is suitable for performance testing of mechanisms. With this in mind, a pin-ondisk tribometer was designed and co nstructed that would be capable of operating in this pressure

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24 regime. Many of the friction studies performed on solid lubricant coatings are run using pin-ondisk tribometers. There are several advantages for this type of tribom eter. The disk is only expected to run in one directi on meaning a simple motor can be used. The friction forces are only measured in one direction so it simplifies th e load sensing requirements, and these systems can typically generate large numbers of cycles in a short period of time. This can be beneficial when trying to run a low-wear film to failure. As mentioned each tribometer is driven by a high vacuum compatible motor capable of operating at pressures below 10-6 Torr. The motor that was chosen was a T-Max 5 servo motor from Nutec Components Inc. ( www.nutec1.com ). The unit is a m otor and stage combined into a single housing. The drive is from a linear moto r that has been wrapped around a central axis giving it the ability to turn the stage. This motor is capable of generating 2.9 N-m of torque continuously. This value is much higher than any torque expected fr om a pin-on-disk type contact under loads of 1 to 10 N. The spindle can be commanded to spin at speeds ranging from 0.001 rpm to 1,000 rpm allowing for a wide range of operating speeds. Nutec also outfits all its motors with positional encoders so the controller can determine the current position at all times. The encoder built into this system has 72,000 lines resulting in a resolution of 18 arc-seconds. The load cell for this tribometer is an AMTI ( www.amtiweb.com ) MC-2.5A six-axis transducer. The choice to use th is lo ad cell was based on availa bility. At this time no other manufacturer could guarantee vacuum compatibility to the required levels. The transducer was modified to ensure vacu um compatibility below 10-7 Torr. The main changes are removing any paint and decals from the device and replacing al l the wiring inside with Teflon coated copper wires to reduce outgassing. This load cell ha s a maximum capacity of 450 N in the axial

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25 direction and 225 N in the radial directions. The re solution of the load cell in these directions is 15 mN meaning the lowest resolvable fricti on reading under a 1 N normal load is 0.015. To manipulate the sample, th e entire assembly is mounted to a set of Schneeberger ( www.schneeberger.com ) m icrometer stages. The purpose of these stages is to accurately adjust the track diameter to fix the sliding speed, and to load the sample against the surface after properly positioning the pin. Figure 2-2 Pin-on-disk design vari ations (a) Initial pin-on-disk design used flexures and a compression spring to apply the load to the system. Also, note the orientation of the load cell. In this configuration its le ast sensitive axis is in the normal loading direction. (b) The final design uses a can tilevered pin holder to apply the load drastically increasing the stiffness of the sy stem. The load cell is now mounted so the normal force and friction force are measured by the most sensitive axes of the transducer. The machined parts in the pin-on-disk assemb ly went through a major revision. Initially, the device was loaded using a flexure and comp ression spring assembly (Figure 2-2a). After construction it was found that the dynamics of this system made it difficult to accurately measure friction, particularly if the su rface had a misalignment. Any misalignment between the surface of the disk and the pin re sulted in a sinusoidal fluctuation in no rmal force. The stiffness of the system was low enough that the vibrations indu ced by this forcing function caused the pin to

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26 bounce off the disk surface. As a result, a new approach was taken in the revision where a cantilevered sample holder was designed and used to apply a load to the disk (Figure 3b). In the new design, the load cell orientati on was altered so the two most sensitive axes were measuring normal load and friction force. The new revisi on simultaneously simplified the design, improved the sensitivity of the measurements and increa sed the stiffness of the system. While any misalignment between the motor and the pin still caused sinusoidal fluctuations in the normal load, the value was accurately measured and there was no evidence that the pin ever left the surface of the disk as it had with the first design. The sample holders for this tribometer are cap able of housing multiple disk sizes as well as multiple pin sizes. The modular disk housing has a base component that bolts to the face of the motor. A second component is screwed down over the disk pinching it agai nst the base plate and aligns the disk face with th e face of the motor. The housing can accommodate samples of varying thickness because of a compressible o-ri ng of Viton behind the sample that keeps it pressed to the alignment surface on the housing. The ball holder is machined into the load flexure for this tribometer. As with the disk ho lders, there are multiple flexures to accommodate different diameter pin samples. The sample is inserted into the holder and a set screw is tightened behind the ball holding it firmly in the housing and preventing it from slipping during the experiment. The acquisition of data is done using a 6036E data acquisition card from National Instruments ( www.ni.com ). Each of the channels (norm al fo rce, friction force, ambient pressure and motor position) are recorded at up to 10 kHz (a value of 1 kHz was typically used). The length of time over which data is collected is variable to allow for different cycle times. A general guideline was to collect data for a minimu m of two cycles to help reduce the effects of

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27 noise in the data as well as any surface anomalies that may skew the result. This data is then averaged and appended to the average data file along with a timestamp. Phase-locked data is also available from this tribometer. This data is periodically collected an d stored in a separate data file for further analysis. 2.4 High Vacuum Linear Reciprocator A primary issue with pin-on-disk testing is contact pressure for this type of a sphere on flat geometry is very high. Usually the contact pres sures are well above the yield strength of the coating materials. If a larger pin is used to reduce contact pressures, th e sliding speed across the contact is varying as a function of radius. To co unteract this effect, the wear tracks are usually narrow and can be difficult to analyze. Many mi croscopy techniques require the area of interest to be between 100 m and several mm for an accurate sampling (e.g. x-ray photoelectron spectroscopy). One alternative tribometer design is a linear reciprocator. This type of tribometer allows for a wide range of pin geometries. These pin ge ometries can greatly reduce the contact stresses on the films and allow a much larger wear track that is more easily analyzed in an XPS or other spectroscopy equipment. Adjusting the surf ace temperature of a sample on the linear reciprocator is also much easier than a pin-on-disk. The stage moves in a linear reciprocation over a fixed distance, so flexible cryogenic lines can be attached to the stage for adjusting the surface temperature. In a pin-on-disk contact, th e disk sample is rotating and creating a viable conduction path for cooling the surface is not as straightforward. The stage used in this design is a Lineax 10 from Nutec Inc. ( http://www.nutec1.com ). The m otor and stage are capable of operat ing at high vacuum levels (pressure < 10-6 Torr). The motor itself is a brushless linear motor and is equipped with a non-contac ting encoder system. The maximum speed of the motor is 3m/sec with a positional reso lution of 0.5m and a

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28 repeatability of 2.5m. The stag e is capable of handling a norm al force of 1000 N and tangential force of 325 N. However, as the load is increased so is the current requir ed to drive the stage, resulting in larger amounts of h eat generated due to resistive h eating in the motor. A closedcircuit cooling system was embedded in the mo tor housing which can be hooked up to a chilled water system to maintain a safe operating temperat ure. This is in contrast to the pin-on-disk tribometer which does not require cooling due to the low forces and torques generated in those experiments. The maximum track length of 50 mm is defined by the motor travel. The mounting plate on the stage has an array of tapped 4-40 holes to allow for a variety of counterface positions. The load cell used in this system is the same type as described earlier for the high vacuum pin-on-disk tribometer, a MC-2.5A from AMTI. The only di fference is the load capacity and resolution of the load cell. This system is de signed to work with loads as high as 1000 N, so a 2200 N load cell is used. With all of the load cells from AMTI (the MC-2.5A included) the resolution in the normal or z direction is twice th at of the x and y directions. To enable small friction coefficients to be measured by this syst em the more sensitive x and y axes are used to sense the friction force. As mentioned, the resolu tion of these load cells is scaled based on their maximum capacity. The transducer used here cannot resolve a friction force less than 75 mN. However, the normal loads intended for the tr ibometer are between 100 and 1000 N resulting in a minimum resolvable friction coefficient of 0.001 at 100 N. Due to the size of the linear stage, several loading options were precluded. Ultimately, the choice was made to use a cantilevered arm load ed with a compression spring, which can provide a variety of loads by varying the stiffness of the compression spring. The system works by using a t-shaped arm that pivots about a pair of beari ngs mounted to the frame of the tribometer. As

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29 the spring is compressed, it imposes a force on the top of the t-shaped component in the horizontal direction. The body pivots about the bearing axis and imposes a vertical force on the stage to balance the moment. There is no mechani cal advantage in this system; theoretically, the load applied by the spring is the same as the one applied to the stage. This was done to make it easier to estimate the required spring stiffn ess and compression for a desired load. One advantage of using a simple design is low num ber of components. Every moving part and surface interface in a vacuum system is a possibl e source of virtual leaks and outgassing. By minimizing the number of mating components, it is possible to improve pump down times and sample throughput. Figure 2-3 Components of the high vacuum linear re ciprocating tribometer. (a) Schematic of the high vacuum linear tribometer (b) Pin sample holder as sembly. (c) The compression spring pulls the t-shaped lever arm forw ard imposing a normal force on the surface between the sample and the counterface. Although the assembly works using a similar theo ry to the original pin-on-disk device, the dynamics of this system should not result in the same issues as the pin-on-disk. This is primarily

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30 due to the higher spring stiffness and large no rmal loads applied to the surface. Another difference between the systems is the forcing func tions. The pin-on-disk tribometer generates a sinusoidal forcing function with a frequency eq ual to the cycle frequency. The linear reciprocator does not generate th e same type of forcing functi on. Any misalignment between the surface and the pin will manifest itself as a ramp with a frequency equal to the cycle frequency. 2.5 Cryogenic Pin-On-Disk Tribometer At the core of the apparatus is a Falex PinOn-Disk tribometer (see figure 2-5). This system uses a dead-weight load to apply a normal force onto a pin sample. The pin remains stationary while the disk below it spins. The shear force acting at the interface between the pin and disk results in a friction force on the pin. Th is force pulls the armature attached to the pin inducing a strain across the load cell connecting the armature to ground. The force is measured by the deflection of a strain gauge and fed into the internal electronics of the tribometer to calculate a friction coefficient. The user inputs the mass of the dead-weight load attached to the armature. The friction coefficient is calculated ba sed on the ratio of the distance from the load cell to the gimbal and the pin sample to the gimbal. A more detailed discussion of the friction calculation and the uncertainty in the measurements from this tribometer can be found in Appendix A.

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31 Figure 2-4 Schematic of modified pin-on-disk tribometer used in the temperature studies. The forces in the system are depicted in the image above using grey arrows. There were two major changes from the as-r eceived Falex system to better accommodate the testing. The first change was the disk holde r. The second generation holder was created for easy sample changes. The housing itself is made up of two parts. The lower half is mounted to the spindle driven by the motor. This portion a ligns based on the spindle shaft and is tightened onto the shaft using a clamping mechanism. The top portion is modular so that multiple sample sizes can be used. Currently, tw o sample sizes can be used. Th e first is a 2.0 diameter disk with a range of thicknesses from 1/8 to 5/16, wh ich was used for all the testing described here. The second is a 1.0 diameter disk with the same thickness range. The lower half of the housing has a reverse-threaded stud coming out of the to p. The upper half threads onto this stud and is tightened using a tool provided by Falex. The ho les on the outside of the upper half are used by the tool to apply more torque for tightening the sample housing. The second major change was the pin holder. The holder was similar to the original Falex holder in that it had a hexagonal head to allow mu ltiple tracks to be run on a single ball. The

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32 difference comes in how the ball is attached to the housing. The large temperature fluctuations in these experiments proved a problem for the orig inal holder because the samples were glued to the housing. It was necessary to change the de sign and create a mechanical connection between the ball and housing. This was accomplished by using a threaded stud that runs down the center of the housing. The balls used must have a 440 threaded hold tapped in them, the stud has 4-40 threads on the lower half and left hand threaded -20 threads on the upper ha lf. Once the ball is threaded onto the lower half, the stud is rotate d back into the housing ca using it to tighten the ball further and provide a strong fo rce to keep the ball in place. Once the sample is mounted, decreasing the temperature causes the ball to be further pulled into the housing because the stud inside is made of aluminum while the housing is stainless steel. Figure 2-5 Cryogenic pin-on-disk sa mple housing. (a) Thread the ball onto the lower shaft using a standard counter-clockwise motion until it is tight. (b) Rotate the upper shaft in the clockwise direction to retract the ball into the housing. The ball will fit snugly into the mating surface inside the pin housing. (c ) Section view of th e assembly after the ball has been properly tightened. The materials for each component are also highlighted to indicate the system will further pull the ball into the housing as the temperature of the system is decreased.

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33 2.6 In Situ Wear Tribometer The ability to determine the life expectancy of a solid lubricant coating is of great importance to designers. Mechanisms must be desi gned to operate within a useful lifetime, but it is difficult to accurately predict that life without some guidance on how the system will fail. Solid lubricant coatings generally fail due to some type of wear mechanism be it delamination, abrasive wear, plowing etc. Char acterizing the type of wear a so lid lubricant coating experiences and its severity can improve designs where th ese materials are the limiting factor. To accomplish this, a tribometer was designed and cons tructed for the purpose of measuring wear of a surface in situ. Wear track topography can vary widely based on geometry of the contact and composition of the surfaces. In the ca se of many solid lubricants wear rates are between 1x10-4 and 1x10-8 mm3/Nm. This means it could take several thousand cycles before a nanometer of the surface is removed. Given the sensitivity required to captur e the surface evolution, the decision was made to use a scanning white light interferometer (SWLI). The Zygo New View 5030 was chosen because of the quality of its opt ics, and a feature height resolution on the order of angstroms. This system is equipped with set of motorized st ages capable of adjusting the sample orientation (roll and pitch) and position (x, y, z) to provi de the best surface scans. The difficulty was designing a tribometer capable of f unctioning in this limited workspace. Like the vacuum tribometer, finding a motor and stage system suitable for the intended environment proved difficult. Initially, a small st age was chosen which was driven by a brushed DC servo motor. The motor was found to have problems running for extended periods in low humidity environments. This is most likely due to the graphite br ushes, commonly found in many servo motors, becoming brittle in an enviro nment that lacks moisture. The second version of the stage was a Parker 401-XR linear stage wi th a HV172 stepper motor. This system is

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34 capable of sustaining a much larger load than th e previous version (200 N compared to 5 N) and came with a linear encoder for monitoring stage position. Measuring abso lute position of the stage provides better repeatability in the m easurement that estimating the position using motor rotations. More specifically, the coupling between the motor and stage has dead zones that cannot be accounted for when the position is bein g estimated by motor revolutions. The stage and stepper motor combination have a positional repeatability of m. The repeatability is important when producing the time-lapse images of the track because any systematic drift of the stage is readily apparent in th e resulting video. The new stag e also expanded the range of possible sliding velocities with a maximum velocity of 50 mm/sec. Figure 2-6 Overview of in situ tribometer. (a) Schematic of in situ tribometer. (b) Illustration of loading mechanism. (c) Drawing of tri bometer placed on top of the Zygo stage. The intended normal load for the tribometer was less than 10 N which required a load cell that could read friction forces as low as 50 mN. A JR3 50M31A load cell was chosen with a resolution of 28 mN in the normal load directi on and 14 mN in the fric tion force direction.

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35 Following the design philosophy mentioned earlier, th e six-axis load cell was placed directly in the load path between the pin sample and ground. This is believed to be the best way to minimize uncertainties in the friction and normal force measurements. The application of load in this tribometer had to be small to allow the system and a surrounding environment chamber to fit inside the framework of the SWLI. The size constraints on this tribometer limited the options for the load application mechanism. The design employs a leaf-type parallelogram flexure attached to a microm eter stage to apply a load to the pin sample. One benefit of this type of design is the wi de variety of normal lo ads that it can apply by changing the geometry of the leaf flexures. A s econd property of this type of flexure system is its high stiffness in the friction force direction. This will help to prevent the system from rotating as the force between the pin and the counterface increases. By adjusting the micrometer stage, the pin sample is brought into contact with the co unterface. The load increases as the deflection in the leaf flexures increases. The six-channel load cell is used to read th e applied normal load in real-time to allow the user to accurately rese t the load after the pin has been brought off the surface for imaging purposes. The pin sample holder is a PEEK component mach ined to hold the ball at a 60 angle with the surface. The purpose for this is to maximize the number of tracks that can be run on a single ball. By angling the ball in this way, a single sphere may be used for up to 6 tests. This matches the number of tracks that can be run on the rect angular coupons. The counterface is mounted to a machined plate attached to the motorized stage. The hole-pattern on the plate is such that the counterface can be held in three different positions. Each position can have two tracks run on it for a total of six tracks per sample.

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36 2.7 Cryogenic High Vacuum Bushing Tribometer Solid lubricants and fluorinated greases are commonly used in bushing configurations to provide continuous operation in extreme environm ents where many traditional lubricants are unable to operate. Thermal limits (high a nd low) often preclude the use of many hydrocarbon oils. High vacuum environments are particularly challenging and there are a number of moving mechanical assemblies that utilize bushings to provide low torque opera tion in space and high altitude vehicles. Bushing contacts are a typically designed to be closely matched axis-symmetric bodies of revolution (shaft and through hole). Such a comm on and practical device, bushings are used in everything from door hinges to jet-engine actuat ors. Surprisingly, there is little published on component level testing of bushings; perhaps, due to the multitude of quiet complexities found in such simple and ubiquitous components. These complexities include e volving geometry during operation ( 45 ),and uncertainties in contact area, pressure distribution, a nd frictional forces. In an effort to study the extended performance of a single bushing component, a cold thermal vacuum bushing tribometer was constructed. The design of the tribometer followed the methodology described by Schmitz et al.( 39), which essentially describes the importance of having the load path flow through a 6-channel load-cell that reacts the normal and frictional fo rces and moments near the point of contact. A six-channel load cell suitable for high-vacuum operation was designed an d fabricated by AMTI (Boston, Mass) and having a maximum load of 1000N in the loading direction and maximum of torque of 50Nm. The load cell is a strain-gauge based six-channel device used to sense the load that is transferred from the flexures into th e contact between the bushi ng and the shaft. It measures the normal load applied to the bushing as well as the frictional to rque generated at the

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37 bushing-shaft interface. The forces and torques are output by the load cell electronics as analog voltages that can be recorded as the test is running. Traditionally, dead weight loads, pneumatics, or hydraulics are selected to apply loads in tribological testing. The large lo ad requirement for this applica tion (500N) precluded the use of dead weight loading systems in the vacuum chamber, and the modular methodology described in the introduction precluded pneumatic and hydrauli c loading/feedthroughs in the system. A relatively simple spring loaded system was select ed to apply a normal force to the sample; such systems are typically avoided because creep and wear in the sample act to reduce the strain in the spring and the normal force is continuously decaying during operation. In this system the continuous measurement of normal force reduces these time varying biases. Additionally, very soft springs with the appropriate load capacities were selected to maximize spring deflections at load and thus minimize variations in load due to wear and other gradual deformations that occur in during testing. Figure 2-7 Bushing tribometer

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38 Due to the challenges associated with maintaining low friction forces in vacuum environments, efforts were made to eliminat e bearings and bushings in the design. The tribometer uses bearings only in the support of the rota ting shaft and the motor. To apply normal load a series of two flexures act as a pivot to transmit the vertical load and as a restraint to force linear motion of the bushing in the loading di rection. The load spring is compressed by tightening a nut on a threaded rod th at is affixed to the load flexure. The alignment flexure is designed such that it can only move in the horizontal direction and imparts a purely horizontal load on the load cell and sample assembly through a point contact with a sphere. The function of the alignment flexure is to accommodate deflecti on during the wear of the bushing. The flexures were wire electro-discharge-machined from bulk monolithic stainless steel pieces. The drive system for the shaft uses a high-vacuum compatible servo motor capable of running at speeds as low as 0.001 rpm and as hi gh as 1000 rpm. The motor was originally designed to be cooled via convection. However, since this is not possible inside the vacuum system a chilled water circulator and aluminum block at the base of the motor are used to conduct heat out of the motor at high-vacuum levels A flexible coupling is used to connect the motor to the rotating shaft and provides the opportunity to easily vary the diameters of the shafts. The shaft is aligned using two high vacuum rolling element bearings lubricated with fluorinated grease. These greases have very high molecu lar weights (3000+ AMU) with vapor pressures that are below 7x10-7 Torr at room temperature. The entire bushing assembly is bolted directly onto to the front of the load cell to reduce uncertainties in the friction coefficient m easurement. As described by Schmitz et al .( 39), multiaxis load cell should be place directly in the load path between the sample and ground as close to the contact as possible. An illust ration of the load path is shown in figure 2-9. All loads reacted

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39 by the shaft are carried through the load cell, which is supported and constrained to linear motion by the alignment flexure assembly. Figure 2-8 Illustration of load path to ground for bushing tribometer design. There are a number of vacuum tribometers; however, the ability to perform cryogenic vacuum testing is much more limited. To reach temperatures below -100C liquid nitrogen was fed through a cold flask within the vacuum cham ber. This flask is mounted approximately 75 mm from the sample and thin copper braids ar e bolted from the reservoir to a copper bushing housing. This short conduction path and high thermal conductivities of copper made rapid cooling (~10C/min) of the bushing assembly possible. To prevent cooling of the load cell, a polyetheretherkeytone (PEEK) insert is bolted between the copper bushing housing and the load cell as an insulator. Th e temperature is read from a thermocouple rigidly held to the outside of the bushing. The temperature rise across the bushing can be estim ated using a simple 1-D heat

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40 transfer analysis. The value for this temper ature rise is 3C using a 100 N normal load, 10 mm/sec sliding velocity and a friction coefficient of 0.5.( 46) Acquiring data during an experiment requires a card capable of reading multiple analog signals simultaneously. Using da ta acquisition all the different outputs (8) are read at up to 10 kHz for variable lengths of time. The ability to adjust acquisition rate and time is useful because the cycle time varies from one test to the next depending on the desired spindle speed. Typical settings for these options are an acquisition rate of 1,000 Hz per ch annel for a time period of six seconds. A majority of the tests run have been at 20 rpm, so the data is acquired for two complete cycles before being processed. All of the data collection is phase locked with the motor position. Plotting the force and torque values with respect to the motors angular coordinate allows identification of persistent features. Any eccentricity between the shaft and bushing will manifest itself as a si nusoidal fluctuation in the normal force when plotted versus the angular position of the shaft. These fluctuations are by design small compared to the applied normal load. This data collection scheme occupies a grea t deal of memory. The approach that is frequently used computes average cycle values from the kHz data and stores the average values in a single file. Periodically, the phase locked da ta is stored in a separate file that is time stamped with the cycle number. The distance va lue is calculated by using the input spindle speed, the shaft radius and the time the test has been running. This value has uncertainty because the values of spindle speed a nd shaft radius are assumed inst ead of monitored throughout the test, however, distance is not a factor in calculating fric tion coefficient. Schmitz et al. performed a rigorous uncertainty analysis on a system using the same electroni cs as the bushing tribometer

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41 and found the contribution of the uncertainty fr om the electronics to be negligible when compared with other error sources.( 39) Computation of Friction Coefficient The strain-gauge based load cell outputs forces and torques about all three axes, however, some of this information is not used in the comput ation of friction coefficien t. In this tribometer, the two force values Fx and Fy are used to calcul ate the total normal force exerted on the shaft. The only torque that is of interest is the one generated by the fricti on between the shaft and bushing, thus only three of the six load cell channels are re ad into the software. The friction coefficient is defined as the dime nsionless ratio of the friction force between two bodies divided by the normal force pressing th em together. The normal load exerted on the shaft by the bushing is measured directly, and the frictional stresses result in a torque. Assuming both the shaft and pin are rigid bodies (illustr ated in the figure 2-10b) the normal load and friction force are assumed to be point loads and the standard friction coe fficient equation, a ratio of the friction force to the normal force, is given by Equation 1. f nsnF T FRF (2-1) For two bodies in contact that are not infinite ly stiff the contact area is finite due to deformation. This deformation introduces an error into the friction coefficient value reported by Eqn. (2-1). A simple model is to assume the sh aft contacts the bushing over the bushings entire length and the contact occurs ove r a specific wrap angle. Without more advanced material properties and contact mechanics, the exact wrap angle and pr essure distribution cannot be determined. However, assuming a uniform pressure distribution over a contact area an analytical solution can be derived. For the cas e where the contact happens from to the normal load

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42 measured by the load cell is the integrated contribution of the pressure distribution illustrated in Figure 2-10c. Figure 2-9 Friction coefficient derivation for bushing tribometer results. The normal load measured by the load cell is a function of the length of the bushing (w), the radius (Rs), the pressure (p) a nd the contact angle ( ). 02cosnsFpwRd (2-2) The average pressure (p) can be calculated as a function of given the measured normal load Fn. 2sinn sF p wR (2-3) The frictional shear stresses are simply the produc t of the friction coefficient () and the normal pressure (p); thus, the frictional torque can by found from the following integral. 2 0' 2 2 sin()sin()nn s s sFF R Tl R d wR (2-4) Following Eqn. (2-1), the computed friction coefficient () is given by Eqn. (2-5). sin() (2-7)

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43 For small wrap angle ( ) the error is likely negligible, and the reported friction coefficient () is always larger than the true value (). A plot of the percen t error (defined as %100' error ) is shown in figure 2-11 illustrating that at a wrap angle of 20 the error in the reading is still below 5%. Figure 2-11 Error in friction coefficient base d on the contact angle between the bushing and shaft.

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44 CHAPTER 3 VARIABLE ENVIRONMENT EXPERIMENTS 3.1 Overview of Materials Used in Experiments This series of experiments was meant to surv ey the field of possibl e coating options to determine their performance under varying environmental conditions. Groups from Harris Corporation, the Air Force Research Laboratorie s (AFRL) and the University of Florida were gathered to give input into the material selecti on for these tests. The coatings ranged from hard metallic coatings used in gears to intricate composite coatings designed specifically for low friction in varying environments. All coatings were applied to 1/4 spheres and 2 diameter disks of aluminum 7075. This substrate wa s chosen because it is common in many space mechanisms due to its high strength-to-weight rati o. The coatings were applied to both the pin and counterface to examine how the materials respond in a self-mated contact condition, as opposed to the coating against a nascent aluminum surface. Table 3-1 List of coatings initially tested for environmental sensitivity

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45 Some of the coatings listed are commonly found in space applica tions, hard anodized aluminum with PTFE, MoS2 with titanium, MoS2 with Sb2O3 and gold, etc. The MoS2 with Sb2O3 and gold coating was originally developed at the Air Force Research Labs, and has become a commercial standard for low friction in terfaces in a variety of mechanisms. While many of the composites listed are commercially av ailable they are not cu rrently used in space applications and some are still in developmental stages (MoS2 with Sb2O3 and graphite and near frictionless carbon). The purpose of this variet y was to give designe rs insight into the performance of these new materials as possible options for future designs. These coatings are created using a variety of different techniques. The anodized coatings are created by soaking the sample in a sulphur ic acid bath while put ting a current through the system. The coating forms about 2m thick, wh ile roughly 1m of the sample material is removed leaving the sample oversized by 1m. The electroless nickel process is a similarly bathed in solution. The system does not require electricity which reportedly reduces the friction coefficient over electroplated Watts nickel. Both of these coatings can be created with PTFE as a solid lubricant on the surface. 3.2 Cryogenic Pin-On-Disk Experiment Environmental Protocols Reducing the temperature of any surface at atmospheric pressure (as opposed to vacuum) can form water on that surface as water vapor in the surrounding environment comes out of solution and condenses. When performing experime nts at reduced temperatures it is important that the test be run above th e dew point for the water vapor in the atmosphere to avoid confounding the tribological results by formi ng water at the interface. As the surface temperatures decrease, the amount of water vapor required for condensation to occur decreases. To combat this problem, it is necessary to remove as much water from the surrounding

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46 environment as possible. The enclosure used in this experiment was a Vacuum Atmospheres environment chamber, which isolated the experi ment from ambient humidity. This chamber was backfilled with ultra-high purity nitrogen gas (99.999% pure) as well as boil-off from liquid nitrogen dewars to remove as much water vapor from the system as possible and create an inert environment with controlled relative humidity levels around the tribometer. The environment chamber was also fit with a humidity sensor by GE Sensing capable of resolving relative humidity less than 0.1%. Inside the chamber, a technique similar to th at used in metal inert gas (MIG) welding was employed to keep the surface of the disk as cold and clean as possible. Liquid nitrogen flows from a pressurized dewar through braided steel line s to a nozzle located a bove the disk surface. The liquid nitrogen flows from the nozzle two inches above the disk surface nearly parallel to the plane of the disk. Another jet of nitrogen from an ultra-high pur ity nitrogen cylinder is pointed down onto the center of the disk. The flow rate of this jet is controlled by an adjustable flow controller to increase or decrease the amount of gas impinging on the surface. The two jets intersect above the disk surface causing the liquid nitrogen to volatilize into cryogenic nitrogen gas which is carried to the surface by the impinging gas nitrogen jet (see Figure 3-1). Using this method, surface temperatures could be controlled dow n to -80C. To increase the temperature of the surface, a heat gun was mounted above the di sk surface impinging direct ly onto the center of the disk. The temperature of the disk surface was increased as high as 180C during the testing. In both cooling and heating, it was important to keep the gas flow near the center of the disk because any viscous flow off-axis to the load ar mature could affect norma l force and or friction force.

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47 Figure 3-1 Description of impinging jet technique used to cool and clean the surface of the disk during experiments. Experimental Procedure The pin-on-disk experiments conducted on the Fa lex instrument were run at the lowest obtainable relative humidity levels (below 1% for all the experi ments). A 500 gram dead-weight load was hung from the armature on the tribometer resulting in a nominal normal load of 2.5 N at the pin. Each sample was tested at three sepa rate temperatures -70C, -30C and 20C. Fresh tracks were used for each test to eliminate the possibility previous temperature affecting the results. The procedure for each new track was to run the sample in at room temperature to wear through any oxide layers that may have fo rmed on the surface between deposition and experiment. After the friction coefficient reached a steady-stat e, the test was stopped and the temperature of the system was reduced to the targ et value. The temperature was held near the target for several minutes prior to beginning th e test to ensure both the disk and ball were equilibrated. Once the system was at the target temperature, the motor was started and friction coefficient recorded for a period of five minutes or more. This was to ensure there were no transient behaviors in the fric tion coefficient and there was e nough data minimize the effects of noise in the system. The sliding speed for the innermost track on the sample was 12.5 mm/sec (5

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48 mm radius at 50 rpm), but the sliding speed increas ed as the track diameter increased because the motor was unable to reliably spin the disk at speeds less than 30 rpm. The maximum sliding speed for the samples at the outermost track wa s 30 mm/sec (10 mm radius at 30 rpm). Johnson et al. showed friction coefficients for MoS2 coatings were relatively insensitive to sliding speed over this range.(10) The steady-state temperature rise in the system can be estimated using a simple 1-D conduction equation.( 46 ) qa T K (3-1) To use this equation, the nomina l contact radius and the heat flux must be calculated. The contact radius (a) can be estimated using a Hert zian contact solution. For a sphere on flat contact, the equivalent radius of the two bodies is equal to the radius of the sphere. 11111abaRRRR (3-2) The composite modulus for the bodies is calculat ed using the modulus and Poissons ratio for both bodies. 2211ab baabEE E EvEv (3-3) Using these values, the contact radius can be calculated given a normal load.( 47) 1 33 4nFR a E (3-4) The heat flux generated is a function of the aver age contact pressure, the friction coefficient and the sliding velocity. qPv (3-5)

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49 In the given experiments, a worstcase scenario for friction coefficients was on the order of 0.5. The highest recorded sliding velocity was 30 mm/sec, and the normal load was 2.5 N. Given all these values, the contact radius and heat flux can be calculated. 0.0515 amm (3-6) 24.5 qWmm (3-7) The resulting steady-state temperature rise fo r these contact conditions is less than 1C (0.966C). In the case of the so lid lubricant films, the friction coefficients were an order of magnitude lower than the example; indicating th e temperature of the system before motion and after motion should be equivalent. After running the experiments at reduced te mperatures, the materials that displayed a strong sensitivity to temperature were tested at an elevated temperature of 180C using the same procedure described above. This was done to eval uate how much of a reduction could be seen at elevated temperatures. 3.3 In Situ Wear Experiment The friction response of MoS2 coatings in dry (RH less than 1%) environments and high vacuum has been well documented. Since the discove ry of this material more than 60 years ago, different techniques have been developed for applying a MoS2 coating, different additives have been mixed with MoS2 to improve its friction response and its surface adhesion. However, the wear analysis of this coating has been very li mited. A majority of the studies where wear volume is quantified report values determined ex situ and usually after the coating has failed. The purpose of this experiment was to capture the evolution of the w ear scar as the test progressed, and determine a correlation betw een friction coefficients and wear rates.

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50 The tribometer used in this experiment wa s a linear reciprocati ng tribometer mounted underneath a Zygo New View 5030 s canning white light interferom eter. Although the resolution of the device is on the order of angstroms, th e device is capable of detecting surface topography changes on the order of nanometers. The genera l idea behind this experiment is to image a section of the wear track at di fferent cycles throughout the test to estimate a volume loss and a wear rate. Sample Description The counterface samples used were nominally 1.5 x 1.0 x 0.1875 coupons made of 304 stainless steel. The coupons were polished on a polishing wheel to a surface roughness less than 50 nm prior to being shipped to the suppliers for coating. Th e pin samples were 6061-T6 aluminum spheres that were drilled and tapped for a 4-40 screw. No surface preparation was done to the spheres prior to coat ing. The coatings applied to the pins and coupons were all MoS2 based, but had a variety of other constituents. The commercially available coatings used were MoS2 with titanium, MoS2 with Sb2O3 and gold, and MoS2 with nickel. The newly developed coatings from the AFRL were MoS2 with Sb2O3 and one of the chameleon coatings (MoS2 with Sb2O3 and graphite). All coatings were nom inally 1m thick, although each one was deposited using different techniques. The MoS2 with titanium is a layered coating created by sputtering a layer of pure titanium on the surface of the substrate, then co-depositing a layer of MoS2 and titanium, then a layer of pure MoS2. The MoS2 with nickel and the MoS2 with Sb2O3 and gold coatings are spu ttered coatings of MoS2 with other constituents to improve adhesion to the substrate, environmental sensitivity and tough ness. The AFRL coatings are created by laser ablating a target made of the desired constituen ts resulting in a coati ng with roughly the same composition as the target.

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51 Environmental Protocols There were two environmental protocols used to test thes e coatings. The first was a cycling environment similar to the pump and purge experiments described in the high vacuum pin-on-disk experiments; in th ese experiments the acrylic chamber surrounding the tribometer was back-filled with dry nitrogen to a relative hu midity less than 2%. The sample was run for a total of 500 cycles in this environment and surf ace scans were taken at 1 20, 30, 40, 50, 100, 200, 300, 400 and 500 cycles. The chamber was th en purged with labora tory air (RH > 20%) and the sample was run for another 500 cycles usi ng the same scanning frequency. This process was repeated once more for a total of 2000 cycles. The hope was the data would provide insight into how the system wears in the transition between a dry environment, where friction coefficients were below 0.05, a nd a humid environment where friction coefficient was above 0.1. Experiments were run starting in nitrogen or humid air to see if the initia l run in affected the future performance of the coating. The second procedure was aimed at determini ng a steady-state wear rate for each coating in a dry nitrogen environment and a humid air e nvironment. Two separate tracks were run on each sample. The first track was run in dry nitrog en at a relative humidity of less than 1%. The second was run in laboratory air at a relative humidity of greater than 20%. Each sample was run for a total of 10,000 cycles and images were taken every 1000 cycles. The chameleon coating from the AFRL was the coating of interest in this experiment primarily because it is hypothesized that the coating draws the favorable so lid lubricant to the surface of the film based on environment. If this theory is correct, the track run in air should have a higher carbon signature than the one run in nitrogen because the gr aphite in the material is designed to be the solid lubricant in the moist environment. Appendix C shows the resu lts of Auger electron spectroscopy run on the Cham eleon coating and the MoS2 with titanium coating.

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52 Experimental Procedure Before any testing could begin with the coated samples, a calibration step was taken to ensure the objective lens of the Zygo was ali gned with the wear track, and to determine the distance between the Zygo lens and the pin hol der. An uncoated counterface and pin were mounted to the system and the counterface was moved under the Zygo objective. The roll and pitch of the Zygo stage was adjusted to align the counterface normal with the objective axis. After the surface was brought into focus, the stage was returned to its initial position and the pin was loaded to 5 N. The sample was then run for 50 cycles on a 5 mm track to ensure a noticeable wear scar was generated. The pi n was unloaded and brought up from the sample surface. The stage was jogged 0.5 mm at a time until the scar was inside the field of view of the Zygo lens. The Zygo stage was adjusted in the x-y plane to bring the end of the scratch into the center of the field of view of the Zygo. Th e distance between the pin and lens was then determined and recorded for future use during the experiment.

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53 Figure 3-2 Screenshot of LabView software wr itten for the in situ wear experiments. When calibration was complete the sample counter face was bolted to the motorized stage. The coated aluminum pin sample was screwed to the PEEK housing and the housing mounted to the load cell. At this point, the LabView software was opened to begin the experiment. Inside the software there are several step s required to begin a test. 1. Enter the calibration constant s for the load cell, thermocouples, humidity sensor and positional encoder. 2. Set the location for data file s to be saved and enter any notes on the pin and counterface samples 3. Tare the load cell

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54 4. Define the absolute motor positions of the Zygo objective lens, the track start point and the track end point so the motor can be commanded back to those positions for the duration of the experiment 5. Move the sample under the Zygo objective and record the cycle 0 image of the track. This image represents the unworn surface t opography and will be used to make all the differential volume loss calculations. 6. Command the sample back to the cycle start position. 7. Bring the pin sample into contact with the surface and adjust the micrometer stage to apply the appropriate load (in this case 5N). 8. Define the track length, sliding velocity and approximate cycle time for the test, 5mm, 10mm/sec and 1.0 seconds respectively. 9. Enter the data acquisition parameters (s ampling period of 1.1 seconds and sampling frequency of 1 kHz) 10. Define how often the software saves a comple te cycle of positional data as opposed to simply an average value for a cycle. 11. Enter the position of the center point of the wear track and the percentage of the wear track to analyze (this is to avoid using the data at the reversal points in the average friction calculations) 12. Begin the test. During these experiments, the LabView software recorded normal force, tangential (friction) force, ambient temperature, relative humidity a nd stage position. As me ntioned, the software periodically records the phase-loc ked data associated with a singl e cycle. Phase-locked data refers to a correlation between th e values recorded, specifically friction force and normal force, and the stage position. This can be most useful for initial run in or the onset of failure when a single cycle can be evaluated to look for a portion of the track that is anomalous. Ultimately the goal of the instrument is to iden tify portions of the track with erratic behavior and be sure to image them to give some insight into the failure mechanisms of the coatings.

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55 Following completion of the test, the samples were removed from the tribometer and vacuum sealed to help protect the surfaces from contamin ation and oxidation. Two of the coatings were evaluated using Auger electron spectroscopy to try and identify the composition of the surfaces after wear occurred. 3.4 Cryogenic High Vacuum Bushing Experiment Environmental Protocols Relating friction results obtained using a tribometer in a laboratory environment to actual frictional losses in a mechanism is not straightforward. Bridging the gap between contrived experiments and real world environments is a pr imary goal for many researchers. To that end, experiments were run using the cryogenic high vacuum bushing tri bometer described in chapter 2.4. The purpose of these tests is to compare re sults obtained at cryogenic temperatures on a pinon-disk tribometer with frictional torques gene rated in a bushing-shaft contact under similar conditions. The simulation of a space environment has a broad definition; in this experiment a high vacuum environment over a range of temperatures are the target conditions. The vacuum level of 1x10-6 Torr was chosen as the required pressure level prior to starting the test. At this pressure, the monolayer formation times are on the order of one second, meaning a single layer of molecules (usually water) will adsorb on a ny available surfaces in roughly one second. The vacuum pumps continue to work throughout th e experiment and can reach levels below 1x10-7 Torr. The point of this pressure range is to ensu re that the shaft is not covered with more than a monolayer of contaminant while it is not in contact. The temperature profile used in these expe riments is a ramp that begins at room temperature (20C) decreases linearly to -60 and then steadily increases back to room temperature. Throughout the course of this te mperature ramp, data points are taken at 20C,

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56 0C, -20C, -40C, -60C, -40C, -20C, 0C and 20C. Although data is collected the entire time, thermal drift in the system can only be el iminated using reversal techniques. The purpose of taking data on the reduction in temperature and the increase in temperature is to ensure that friction response is due to a temperature effect and not to unrecoverable damage to the coatings. Material Preparation The bushings used for these experiments were 440C stainless stee l with a 10.0 (+0.05/0.00) mm inside diameter. The shafts were made of custom 455 stainless steel; they were 8 long and had a 10.0 (+0.00/-0.05) mm outer diamet er. The nominal clearance between the shaft and bushing was 0.05 mm in all cases. All the shafts were polished to a surface roughness of less than 100 nm prior to being sent out for coa ting. Both the bushings and shafts were coated with the same material to mimic the self-mated contacts of the pin-on-disk experiments. Only a subset of the MoS2 coatings were chosen for the bushing experiments mainly to verify the temperature sensitivity of the coatings in an application specific contact geometry. The most sensitive coating (MoS2, Sb2O3 and gold), the least sensitive coating (MoS2 and nickel) and an intermediate coating (MoS2 and titanium) were used to coat bushings and shafts. All three of the coatings are commercially available and co mmonly found in high vacuum environments. Experimental Procedure The process followed for the cryogenic high vacuum bushing experiments was outlined by the Harris Corporation. This protocol was also used in the fluorinated grease experiment mentioned earlier. The procedure for each test was: 1. Open the high vacuum chamber. 2. Load the shaft into the alignment bearings 3. Couple the shaft to the vacuum motor 4. Attach the bushing housing to the load cell

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57 5. Tare the load cell 6. Insert the bushing into the housing 7. Open the LabView software written for the experiment 8. Enter the calibration constants for the load cell, motor encoder, pressure gauge and thermocouples 9. Define the location the data files are to be saved 10. Apply the normal load to the bushing 11. Close the chamber 12. Evacuate the chamber to high vacuum 13. At the desired pressure, run 5 revolutions in the forward direction a nd 5 revolutions in the reverse direction. The average value determines any misalignment or drift in the torque cell. 14. Begin the test. 15. At each temperature run the bushing in a clockwise direction to steady-state and then reverse the motor direction and again run to steady state. This allows for an accurate assessment of frictional torque and reduces the influence of drift. The bushing was run to a steady-state friction value at room temperature. Once a friction coefficient was determined at room temperature, the motor was stopped. The liquid nitrogen was turned on and the indirect cooling system be gan to reduce the temperature of the bushing housing and the bushing itself. A thermocouple pl aced against the outside edge of the bushing was used to estimate the temperature in the cont act. After the temperature on the thermocouple reached the desired running temperature, the mo tor was started. The friction was run into a steady-state value while the temperature was held relatively constant (fluctuations on the order of 5C). The temperature profile for the system was 20C, -60C, -20C, 0C, 20C.

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58 CHAPTER 4 RESULTS 4.1 Cryogenic Pin-On-Disk Experiment Determining a Friction Coefficient The calculation of a friction coefficient va lue for each temperature was accomplished by determining the steady-state value of the friction coefficient data over a period of time where the temperature of the system had reached equilibriu m and remained constant. The purpose of this technique was to eliminate transients in the fr iction response due to fluctuating temperatures. The mean and standard deviation of the friction coe fficient values were calc ulated for all the data collected in this period to quan tify the consistency of the friction. For cases where the surface was plastically deformed and/or large amount of de bris were generated, the spread on the data tended to be large compared to the value itself. The solid lubricants, on the other hand, tended to have a consistent friction response with small stan dard deviations given a constant temperature. Experimental Results The thermal response of materials is a prim ary concern for designers of equipment for space applications mainly because the operating ra nge for some of these mechanisms is -100C to 200C. This large range of operating temper atures causes engineers to consider thermal effects that may not ordinarily be problematic. On e such issue is the mismatch of coefficient of thermal expansion between dissimilar metals. This mismatch can distort the geometries when the system experiences temperature fluctuati ons on the order of seve ral hundred degrees. Another commonly overlooked property is the friction response of a material as the temperature fluctuates. Many designers view friction coefficient of a material as a property of the material and not a value that is strongly influenced by geometry and environment. These experiments began with an experiment performed at the Univ ersity of Florida using thin PTFE composite

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59 coatings indicating a sensitivity of friction to temperature.(6 ) While this result is not widely accepted, and others have published data to the cont rary, it lead to the idea that other materials might also display such a response.(48) The first coatings evaluated in this expe riment were hard metallic coatings. The temperature of the samples was decreased from room temperature down to 200K (-73C). An example of the friction result is plotted in figure 4-1. This is the result for the electroless nickel coating, and indicates the degree of scatter in the data. Each data point on the plot represents the average value for multiple cycles of data. These re sults are representative of all the hard metallic coatings tested, and there was no indication that temperature affected the friction response. Figure 4-1 Friction plot of the elec troless nickel coating at -25C. The plot indicates the large fluctuations in friction coe fficient throughout the experiment. The average value for this data was = 0.8 with a standard deviat ion of 0.1. The temperature of the system is also plotted to indicate the small fl uctuations in temperature over a 15 minute period. The erratic behavior of the friction of each of the metallic coatings coupled with small changes in the friction coefficient over the range of temperatures te sted lead to the decision that no higher temperature testing was warranted for these coatings. There was no indication that the

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60 plastic deformation or abrasive wear of metallic coatings responded to temperature fluctuations. While this result is not particularly interesting, it provided an important null result; which was the tribometer itself had no inhere nt bias resulting from a varying temperature. This can be seen from the fact that although the temperature of the system changed by 100C, the responses from each of the metallic coatings stayed constant and those values were unique to each coating. Figure 4-2 Friction response of metallic coatings to varying temperature. The friction forces were scattered due to the large amount of surface deformation and wear that occurred during the test. While metallic coatings are commonly used in gear teeth because they are known to be hard and tough, they are also known to have high friction. One method for improving the performance of these coatings is to deposit a solid lubricant with the coating. Two of the abovementioned metallic coatings are also available with PTFE impregnation. The hard anodize and electroless nickel coatings are commercially available with PTFE. The manufacturers claim these coatings are highly wear resistant, and th e addition of a solid lubr icant greatly improves the friction performance of the coating. These coatin gs were tested following the same procedure as the metallic hard coatings to explore the thermal sensitivity. Like the metallic coatings, there was little evidence to support a th ermal sensitivity in the results from these coatings. There is,

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61 however, evidence to support the addition of PTFE greatly decreases the friction coefficients of these coatings. Figure 4-3 Friction response of metallic coatings with solid lubricant impregnated. (a) The friction results for hard anodize with PTFE coating at -30C. (b) The friction results for the metallic coatings impregnated with PTFE were much more uniform than the metallic coatings alone, but the coatings di d not appear to be sensitive to changing temperatures. Bulk polymeric components can also be used in mechanisms designed for high vacuum such as bushings, snaps and pins. These materials tend to be very inert and have low outgassing rates. The bulk polymeric materials tested in th ese experiments were PTFE, UHMWPE and a PEEK/PTFE. Unlike the coatings tested in thes e experiments, these samples were run against

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62 stainless steel pins. Although the contact does not start out self-mated, polymeric samples readily form transfer films on the steel surface so after a short run-in pe riod they are essentially running in a self-mated configuration.( 49) Figure 4-4 Results for bulk polymeric samples ag ainst a stainless steel pin. The PTFE-based samples displayed a linear decrease in friction coefficient with increasing temperature, while the UHMWPE sample maintained a constant friction coefficient regardless of temperature. The PTFE and PTFE/PEEK/MoS2 composites both demonstrated a linear trend of decreasing friction with increasing temperature. Unfortunately, there are so many properties of PTFE that change with temperature; it is possible to generate a number of explanations for this result. The UHMWPE, on the other hand, did no t show a discernable tr end of friction with temperature. As with the metallic coatings, this is further proof that the system is not inherently sensitive to temperature. In contrast to all the previous results, many of the MoS2 based coatings were very sensitive to temperature fluctuations. MoS2 with nickel was the least sens itive, and reacted similarly to the PTFE based composites. The trend for this co ating was nearly linear with temperature. The

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63 MoS2 with titanium and the MoS2 with Sb2O3 both had a friction coeffi cient of 0.05 at room temperature and increased by 300% as the temperature dropped to -70C. The MoS2 with Sb2O3 and graphite had a friction coefficient below 0.03 at room temperature, but increased to over 0.15 at -70C. In the most extreme case, that of MoS2 with Sb2O3 and gold, the friction increased by an order of magnitude from 0.02 to 0.2 over the 100C temperature drop. Figure 4-5 Friction coefficients of MoS2 based solid lubricant coatings at varying temperatures. The nearly pure MoS2 with nickel coating (95% MoS2, 5% Ni) had the lowest sensitivity to temperature. The commercially available MoS2 with Sb2O3 and gold had the lowest recorded friction at 180 C and the highest increase in friction as temperature decreased (over an order of magnitude). One observation of note was that the materials with the lowest friction coefficient, and highest thermal sensitivities did not have noticeabl e wear scars. This is in contrast to the metallic coatings that exhibited high friction coefficients, no thermal sensitivity and large amounts of debris generation. The differences in wear debris and wear track generation sparked an interest in studying the e volution of the wear scars in situ

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64 4.2 In Situ Wear Experiment The in situ experiments performed in a changing environment provided some insight into the surface topography changes of MoS2 based systems that accompanied changes in humidity. While an increase in humidity immediately triggered an increase in friction, the surface topography did not react as quickly. In fact, th e wear mechanism described by Spalvins where the columnar growths within the film shear and break leaving a well-adhered, thin layer of composite seemed to be accurate for the MoS2 and nickel coating.(21) The other composite coatings where the structure is intentionally amorphous seemed to deform initially under loading and shear stress, but stabilize quickly to a low w ear configuration. Another attribute of the composite coatings is the presence of additive ma terials. These other constituents (e.g. titanium or Sb2O3) are thought to improve the quality and toughness of the coating and prevent the columnar growth of the film. Estimating the wear volume given the topographi cal information of a section of the wear scar required a technique initially introduced by Williamson and Hunt, and refined by Sayles. In the original publication from Williamson and Hunt, the technique was used to evaluate the persistence of asperities after plastic deformation had occurred.( 50) Sayles expanded the technique to evaluate a surface before and after plastic deformation.( 51) The method uses an initial surface topography scan as the basis for the wear scar and takes subsequent topography scans to calculate the cross secti onal area of the wear s car. Using this area extrapolated over the length of the scratch, a volume loss can be estimated. This method was followed for all the coatings tested in both the alternating environm ent experiments and the steady-state wear rate testing. The wear rates reported in this section were obtained using a methodology originally developed by mathematician Stanislav Ulam for predicting odds for the appearance of various

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65 cards in games of solitaire.( 52) The technique was named Mont e Carlo simulation for Ulams uncle, who was a known gambler. The Monte Carlo simulation style is suite d to systems that are dominated by random events or where no analytic al solution can be found. The technique uses random number generation to predict how a random system will react to different inputs. The reaction is simulated a large number of time and th e statistics of the results are considered the odds of one result or another. For the wear rate data each data point on the graph represents an interrupted measurement. Using an estimation of the uncertainty in this measurement and random numbers, new data points were generated. For each data set, 1,000 possible sets were generated and a linear fit was used to calculate a wear rate for each The average wear rate value was the wear rate calculated from the original da ta points, and the uncertainty in that wear rate was the standard deviation of the 1,000 wear ra tes calculated from the generated data. This procedure was followed for all the wear rates calculated in these experiments. Figure 4-6 Methodology for calcula ting wear volume. (a) Plot of a line scan taken from the undeformed surface scan made by the Zygo prior to beginning the experiment and a line scan taken from the surface scan made after 6,000 cycles. (b) Estimation of the wear scar cross-sectional ar ea using the difference between the two line scans. This area is multiplied by the length of the track to estimate a wear volume.

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66 By alternating the environments it was possibl e to determine if the wear rate of the coatings reacted to the environment as the frictio n coefficient obviously changed. In a majority of the coatings, it was found that even though fric tion coefficient changed dramatically, the wear rate of the systems was not affected. The only coating that demonstrated a noticeable and repeatable change in the wear rate with environmental changes was the MoS2/titanium coating. This coating showed a lower wear rate in dry nitrogen when compared with humid air. The coating also wore more readily when transi tioning from a dry environment to a humid environment. Figure 4-7 Plot of volume lost vs. work input into the system. The plot indicates the wear of this coating is sensitive to the partial pressure of water in the environment. In this case, the wear rate initially in air is much less severe than transitioning from a dry environment to a humid one. Many of the other coatings did not show appreciable change s in wear rate in varying environments. For example, the MoS2 with Sb2O3 and gold showed no evidence of different wear regimes throughout the 2,000 cycle test regardless of environment. This sample had a wear rate of 5x10-8 mm3/Nm, but the uncertainty in this value was nearly 100 percent. The deepest penetration depth on this coating throughout the test was 80nm, and the repeatability of the instrument is estimated at 10nm. One explanat ion for the large uncertainty is the methodology used for estimating wear volume was not sensi tive enough to reliably capture surface topography

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67 changes that are on the order of 5nm. To increa se the wear volume the number of cycles was increased to 10,000 in each environment and the nor mal load was increased to 5N. This allowed for the calculation of a steady state wear rate in humid air and dry nitr ogen to discern if the environment had any effect. Figure 4-8 Plot of volume lost vs. work for the MoS2 with Sb2O3 and gold coating. The uncertainty in this value is nearly 100% of the value due to the small volume loss over the course of the entire 2,000 cycles. This result led to testing at higher normal loads and larger numbers of cycles. The 10,000 cycle experiments were aimed at dete rmining a steady-state wear rate in humid air and dry nitrogen. The hope was to determine if environment had any effect on the generation of wear debris. In a graphite system, the lack of humidity causes the graphite to become brittle and wear more rapidly. It was expected that MoS2 based systems would demonstrate a similar behavior when exposed to a humid environment. This hypothesis proved to be correct in every case where MoS2 was the only solid lubricant present in th e system. The wear rates for all these coatings were lower in dry nitrogen th an it was in humid air. Only the MoS2 with Sb2O3 and graphite demonstrated a consistent wear rate regardless of the environment. This coating is also the only one with a const ituent solid lubricant suitable for bot h environments. The theory that graphite is drawn to the surf ace during sliding in humid enviro nments is supported by the Auger analysis in Appendix C. The signature from MoS2 in the track was much stronger in the dry

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68 nitrogen track than it was in the humid air track. Only carbon and oxygen had significant peaks in the wear track run in humid air. The result s for all the steady-state wear studies are shown (Table 4-1). The uncertainty va lues listed are the values fr om the Monte Carlo simulations; however, based on the sensitivity of the instrume nt the minimum uncertainty that should be expected is 1x10-7 mm3/Nm. Table 4-1 All the steady state wear te sting in humid air and dry nitrogen. 4.3 Cryogenic High Vacuum Bushing Experiment Preliminary Experiment To demonstrate the capabilities of the cold thermal vacuum bushing tribometer, an experiment of steel on steel with a vacuum compatible fluorinated grease lubricant was performed. The normal was 100 N leading to a nominal contact pressure of 1 MPa ( 2nsFRl ) at a sliding velocity of 10 mm/sec. Friction measurements were taken over a range of temperatures (20C to -75C) and plotted to reveal th e influence of temperature on the friction coefficient of grease. Reversal techniques are used to eliminate biases in the moment zero. For each data point on the plot, the comput ed friction coefficient was taken after a sliding distance of two meters. The confidence interval s shown on the plot are an indication of the standard deviation of the phase locked friction coefficient data collected at each temperature.

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69 Figure 4-9 Initial experiment using fluorinated grease in a high vacuum cryogenic friction experiment. Solid Lubricant Coating Results Running an experiment in an application spec ific geometry under actu al loads and sliding speeds was the initial intent of this project. In these experiments, three coatings were chosen representing the most temperature sensitive, the least temperature sensitive and an intermediate sensitivity. The tests were begun at room temp erature and the temperature was ramped down to -60C and back to room temperature th roughout the course of the experiment.

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70 Figure 4-10 Plot of the friction resp onse for all three coatings tested in the experiment. None of the coatings demonstrated an increase in fr iction with decreasing temperature in this geometry. The highlighted outlier in the MoS2 with titanium data set was the last data point and could indicate the onset of failure for the coating. Many of these coatings failed after only a few hundred to one-thousand cycles. The plot shows the friction of all the coatings remained low and constant for the length of the experiments. The friction coefficient values for all the coatings are consistent with values found in the cryogenic pin-on-disk experiments. One explanation for the flat friction response could be high wear rates in the bushing shaft geom etry. In the linear reciprocation experiments, the coatings lasted over 10,000 cycles in all cases However, in a bushing-shaft configuration the coating life was often shorter th an 2,000 cycles indicating this geometry may be more severe tribologically. There is some discussion over whether the shaft is riding on an edge of the bushing or if the system is very axis-symmetric The wear rate can be estimated using both assumptions to determine a possible range for the system. In an ideally aligned system where the shaft contacts the bushing across th e entire length of th e bushing, the wear rate is estimated at 5x10-6 mm3/Nm. If the system is running on an edge and only a small portion of the shaft is in contact, the wear rate of the system is 1.55x10-8 mm3/Nm. This range of values encompasses most of the data from the in situ wear experi ments; however, it is not possible to determine if debris generated inside the contac t are playing a role in lubricating the contact prior to failure.

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71 CHAPTER 5 DISCUSSION The term activation energy was introduced by Svante Arrhenius in 1889 to define the potential energy barrier that had to be overcome before a chemical reaction could take place. While the potential energy barrier for a reaction does not change as a function of temperature, the distribution of energy states that each atom or molecule may widen with increasing temperature. With higher temperatures, it is in creasingly likely to find an atom with suitable energy to overcome the barrier and reach a more stable configuration. A schematic of the Boltzmann distribution for three temper atures is shown in figure 5-1. This helps to illustrate the effect temperature has on the system and explains why activated processes occur more readily at higher temperatures. The example of a chemical reaction is typical when discussing thermal activation, but there are other phenomena that demonstrate temp erature sensitivity. Some common examples are viscoelastic creep, grain grow th and diffusion of an impurity. The last of these, diffusion of an impurity (i nterstitial atom) through the latti ce structure of another material is a well documented thermally activated process. It occurs because the atoms are vibrating at very high frequencies 1013 Hz or higher, and an interstitial atom will make a jump from one site to another if the orientation of the surrounding atom s is such that it can make the transition based on its kinetic energy. The amplitude of the vi brations of surrounding atoms, as well as the kinetic energy of the interstitial atom, are domin ated by the temperature of the system. As the temperature of the system increases the oscillations of lattice atoms takes them further from the average position allowing an interstitial atom to squeeze through more easily This is discussed

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72 in greater detail by Porter and Easterling in Phase Transitions in Metals and Alloys.( 53 ) Figure 5-1 Illustration of a potential energy barrier typically found in activated chemical processes. The image to the left of the energy curve is a representation of Boltzmann energy distribution for a large number of atom s or molecules in a system at different temperatures. The portion highlighted in black is the area of the curve containing atoms with energies high enough to ov ercome the potential energy barrier. As mentioned before, a previous publication on the friction response of PTFE coatings initiated these tests. In those experiments, an exponential was used to fit th e results. The results in these experiments seemed to follow a similar trend. The Arrhenius equation was used to attempt to fit the data. The general form of this equation is shown in (5-1). aE RTkAe (5-1) In this equation, k is the rate coefficient, A is a constant, Ea is the activation energy, R is the universal gas constant and T is the temperature in Kelvin. The friction response of the coatings was modeled using this equation in the form shown in (5-2). 011 0aE RTTe (5-2) Here, is the expected friction coefficient at temperature T, 0 is the room temperature friction value, the activation energy is represented by Ea, R is the universal gas constant and T0 is room

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73 temperature in Kelvin. In order to isolate activation energy from the other variable (room temperature friction coefficient), the friction coe fficients are normalized by the room temperature value. 011 0aE RTTe (5-3) A logarithmic transformation is used to create a linear equation in terms of activation energy. This equation can be fit using a least squares regression analysis. 0011 lnaE R TT (5-4) Figure 5-2 Illustration of methodology for creat ing a Monte Carlo simulation of activation energies. The plot of the raw friction data for MoS2 with Sb2O3 and the transformed data can be seen in figure 5-2. Because there are only four data points for each curve, it is difficult to determine an uncertainty in the fit directly. The line represen ted in figure 5-3(a) and 5-3(b) was determined using a linear regression techniqu e combined with a Monte Carlo style of simulation. This style of analysis uses the mean values for each data po int ( and T), shown as circles in figure 5-3, as

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74 the central value in the simulation and generate s a number of alternativ e values, in this case 1,000, that could also be reasona ble given the standard deviati on calculated from the raw data. Figure 5-3 One data series in sta ndard and logarithmic plot with f it. (a) Example of raw friction data with fit of the data. (b) Transformed data with plot of fit. This format is used to generate a fit using l east squares regression. Figure 5-4 Activation ener gy fits for all the MoS2-based composite coatings. This value can be thought of as a thermal sensitivity for the coating. For each of the alternative data series, th e room temperature friction coefficient and activation energy are calculated. The mean value of all the fits is the reported activation energy,

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75 and the standard deviation of these values is repo rted as the uncertainty in activation energy. In many cases, the spread of the data was sizeable, but the resulting uncertainty in the fit was not. Table 5-1 Activation energy values from Monte Carlo simulations of all the data along with the uncertainties in those values. The friction results for the MoS2-based coatings in a pin-on-disk configuration show something about these systems is thermally activat ed. The increase in friction coefficient with decreasing temperature varied based on the coat ing composition. Initially, the hypothesis was that increasing the amount of MoS2 in the coating was decreasing the temperature sensitivity. However, after testing several compositions ranging from 50% MoS2 to 95% MoS2, there was not a strong correlation. As mentioned before, there was a noticeable di fference in the depth of the wear scars and the appearance of wear debris from one coating to the next. The wear rate values determined in the in situ wear experiments offered on plausible e xplanation for the activation energy. By plotting the of activation energy vs. wear rate for each of the co atings, a correlation between the two was revealed.

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76 Figure 5-5 Plot of activation energies determined from the cryogenic testing of all the MoS2 coatings as a function of wear rate determined by the in situ experiments for the same coatings. The MoS2 with nickel had the highest wear rate of all the MoS2 composite coatings. The wear scar on this film grew beyond the field of view of the SWLI af ter 7,000 sliding cycles meaning that wear volumes could not be calculat ed beyond that point. This sample also had the lowest sensitivity to temperature (1.93 kJ/mol). The MoS2 with Sb2O3 and gold wore only one hundred nanometers over the course of 10,000 cycles leading to the lowest wear rate of all the coatings tested. This coating also had the hi ghest activation energy of all the coatings (9.5 kJ/mol). As indicated in the plot, there is a trend of decreasing wear rate with increasing activation energy. The link between these two properties seems to be interfacial sliding. During interfacial sliding, the interac tion between surfaces is dominated by van der Waals forces. In the MoS2 films, the lamellar structure results in planes of the molecule that form sheets with strong bonds. Between these sheets the sulphur atoms interact creating a weak van der Waals bond

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77 resulting in low shear strength for the material. The commonly used analogy for this coating is the deck of cards explanation. The layers of MoS2 act like a deck of cards which can be sheared easily. Figure 5-6 The deck of cards illustra tion for sliding between lamellar MoS2 sheets. The interfaces that are shearing change throughout the sliding event, but the shear force does not change. In this low shear strength configuration, the only interactions between sheets of MoS2 are van der Waals type interactions. It is hypot hesized that van der Waals interactions are temperature sensitive due to a bl urring of the potential surface as the temperature of the system increases. This occurs in a manner similar to that outlined in figure 5-1. Another interesting resu lt was the increase in wear rate as the MoS2/titanium system was cycled from a dry nitrogen environment to a humid air environment. This is perhaps due to a large area of the surface covered with nascent MoS2 that formed during sliding in a dry environment. This nascent MoS2 likely oxidized when oxygen wa s reintroduced to the system causing a large amount of debris generation and an increase in wear volume. The reaction of MoS2 with O2 is a spontaneous one defined by the following equation. 22322724 M oSOMoOSO (5-5) This reaction has a negative change its Gibbs fr ee energy of 2,249 kJ/mol at room temperature (calculations for all chemical r eactions can be found in Appendix D). Initially, the theory was the water in the system caused the degradation; however, fr ee energy calculations found in appendix D indicate water will not spontaneously react with MoS2. The untested surfaces are

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78 likely covered with oxides formed after the sample was created, so running in a humid environment would not cause such a dramatic incr ease in the wear rate. By sliding in a dry environment all the oxides formed on the surface are removed initially le aving a low friction low wear interface to accommodate the re maining cycles. The uncovered MoS2 would be much more reactive to oxygen causing the surf ace to oxidize and degrade quickly.

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79 CHAPTER 6 CONCLUSIONS Three high-vacuum tribometers were designed a nd constructed to test samples in a variety of environments and contact geometries. The variable temperature pinon-disk experiments indicate d an increase in friction coefficient with decreasing temperature. There was a strong correlation between activation energy a nd wear rates for the MoS2based coatings. The cryogenic bushing tribometer experiment support a hypothesis that if the tribological contact is dominated by surface deformation and wear, there will be no sensitivity to temperature variations. The w ear rate in this system cannot be directly measured, but because the system fails after roughly 1,000 cycles, it is a high-wear system.

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80 CHAPTER 7 FUTURE DIRECTIONS Cryogenic high vacuum linear reci procating tribometer experiments are in the immediate future for this project. As mentioned in chapter 2, this tribometer is the easiest to fit with cryogenic cooling capabilities and resistive heating capabilities; making it an obvious choice for continuing to evaluate the temperature response of materials. This system has the best chance of answering the question of why a pin-on-disk experiment in dry nitrogen displayed a strong temperature dependence and a high vacuum bushing geometry did not. Wear measurements can be made for this system using the SWLI before the experiment and after the experiment giving a clear value for volume loss. The in situ wear experiments proved helpful at determining the behavior of these systems to environmental changes. There are also many improvements to the technique that can be made to increase the quality and consistency of the data collected. The first step that can be taken in future experiments is to have the SWLI take several surface scans of an image and average over them to improve the repeatability of the scans. Using a single scan, th e repeatability of the measurements was 40nm. The uncertainty of th e scans is reduced to 10nm by averaging five surface scans. This improvement helps eliminat e surface noise in the sy stem which is readily apparent in the line scans used to calculate wear volumes. A second area for improvement is the calculation of wear volumes. Extr apolating a single line over the entire length of the wear scar can skew the results, particularly if a piece of debris falls into the area where the line scan is being taken. Finding a wear volume based on th e entire surface scan (usually several hundred microns in both directions) can give a better representation of the scar. One method for calculating this value is to use the entire init ial surface scan as the undeformed surface and find

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81 the difference between that and a su bsequent scan. This should help to eliminate the influence of noise in the scan and debris on the surface. Another direction is in the us e of molecular dynamics simulati ons to evaluate the potential surfaces for model system at different temperatures This analysis could provide insight as to why temperature effects are very dramatic in th e low temperature regimes and tend to fall off as the temperature is increased. One theory is the corrugation of the potential surface becomes blurred as the temperature increases reducing the driving forces necessary to cause shear in the solid lubricant interface. Careful evaluation of the potential surfaces at varying temperatures could shed light on this theory.

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82 APPENDIX A UNCERTAINTY ANALYSIS OF FALEX POD TRIBOMETER The uncertainty of the friction coefficients measured by the Falex tribometer can be quantified using the law of propagation of uncer tainty. In this case, there are several contributing factors in the uncertainty of the friction coefficient. They are: Mass of the dead weight load (m) Ratio of lengths from the dead weight load to the gimbal and the pin contact to the gimbal (R2). Ratio of lengths from the friction transducer to the gimbal and the pin to the gimbal (R3). Friction force measured by the strain gauge (Fs). Figure A-1 Free body diagram of Falex pin-on-disk tribometer with friction coefficient calculations. The normal load applied to the sample is cal culated based on the ratio R2 and the mass of the dead weight load. n 2mgmg F~ R2 The law of propagation of uncerta inty indicates the uncertainty of this measurement to be:

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83 2 2 222 nn n2 2 22 222 n2 2 22 22 2 22 4 n nFF uFumuR mR gm g uFum uR RR 9.81 0.59.81 uF 510 0.05 2.0 4.0 uF0.062 The uncertainty of the friction force can also be calculated using th e law of propagation of uncertainty: s3 f 2 22 2 2222 fnn fs23 s23 222 2222 3s3s fs23 2 222 222 2222 f fFR F R FFF uFuFuRuR FRR RF RF uFuF uRuR RRR 1.33 0.051.33 0.05 uF 0.01 0.05 0.05 2.0 4.0 2.0 uF0 .011 The friction coefficient uncertainty value can no w be calculated using the uncertainties of its constituents. f n 22 222 fn fn 22 222 f fn 2 nn 22 222F F uuFuF FF F 1 uuFuF FF 1 0.05 u 0.011 0.062 2.5 6.25 u0.007 Although the calculated uncertainty of the system is 30% of the measured value at = 0.02, the electronics used to output the data to the data acquisition system report friction coefficient as a voltage. The resolution of the sy stem is 1 mV corresponding to a friction coefficient of 0.001; a conservative estimation of the uncertainty of th e electronics is 10 mV corresponding to a friction

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84 coefficient of 0.01. This value will be used as the uncertainty in the measurements for all the data in the pin on disk experiments.

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85 APPENDIX B HIGH VACUUM PIN-ON-DISK EXPERIMENT Environmental Protocols The high vacuum experiments were meant to ev aluate the friction response of the coatings when run in ambient air (relative humidity ~ 40%) and high vacuum (pressure < 1x10-6 Torr). The first test procedure was to run one track from each sample in ambient air then a second track at high vacuum. The load used in these tests was 5 N. This value was chosen so coatings with a low friction coefficient (below 0.05) would still generate a friction force large enough to provide reasonable certainty in the measurement. These experiments were all run at ambient temperature (T ~ 23C). The sliding speed for these experi ments was approximately 20 mm/sec regardless of track diameter. This was made possible by the precise spindle control of high vacuum motor. A second procedure was also used with the hi gh-vacuum pin-on-disk testing referred to as pump and purge. In these tests, the sample was mounted in the pin-on-disk, the chamber was evacuated to high vacuum and the motor was starte d. After the sample ran in to steady-state in high vacuum, the chamber was opened to ambien t air at 40% relative humidity. Once the pressure equilibrated the system was allowed to run to a steady-state condition before the chamber was evacuated again. These tests were performed to see if cycling environments had any effect on the friction response of the coatings. Experimental Procedure Although the high vacuum pin-on-disk tribometer has multiple sample holders to allow for a variety of disk diameters, all the samples used in this experi ment are 2 diameter thick disks and diameter balls. The disk sample was placed on the sample mount bolted to the face of the motor. The cover plate was screwed down over the disk pinching it in place. This design holds the disk flush to the surface and prevents it from moving during the test. As with the disk holders, there are multiple pin flexures to accommodate different diameter pin samples. The sample is inserted into the holder and a set scre w is tightened behind the ball holding it firmly in the housing and preventing it from slipping dur ing the experiment. The sample flexure is attached to the load cell assembl y, and the load cell is tared. This ensures all forces exerted on the pin sample are accurately read by the trans ducer. LabView is started and the program written for this tribometer is executed. The calibration constants for the load cell, pressure gauge and thermocouples are entered into the software. The location where the data files are to be saved is chosen and the program waits for the user to be gin data acquisition. The micrometer stage the where the load cell assembly is attached is adjusted to the desired track radius for the experiment. Once positioned, a second micrometer stage is adjusted to bring the sample into contact and apply a normal load. The normal fo rce is read in real-t ime within the LabView software. When the appropriate load has been applied, the chamber is closed and the system pumped down to the target vacuum level for the experiment. The data acquisition is started in the LabView software which subsequently commands the motor to begin. Average values for ambient pressure, normal force and friction force are recorded throughout the experiment and appended to an average data file for post proc essing. Periodically, one complete revolution of data is recorded.

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86 Humid Air and High Vacuum Testing The sensitivity of MoS2 based coatings to humidity is well documented. Water contaminates the lamellar structure of MoS2 increasing the stress required to shear the material. In a solid lubricant coating, this contamination resu lts in an increased friction coefficient. The expectation for the high vacuum pinon-disk experiments was that the MoS2 based coatings would be extremely sensitive to the partial pressure of water in the system. In metallic coatings, the friction force tends to be dominated by abrasi ve wear and plastic deformation; there is no indication that the presence of humidity woul d improve the ability of these materials to mechanically deform. The polymeric systems, bei ng chemically inert, were also expected to be insensitive to humidity. The table below shows all the pin-on-disk results for high vacuum and humid air. Table B-1 Summary of results from high vacuum pin-on-disk tribometer in humid laboratory air and high vacuum. All tests were run at a normal load of 5N and a sliding speed of 20 mm/sec. As expected, the metallic coatings did not display much sensitivity to humidity in the environment. In all cases, the metallic coatings showed little change co mpared with the scatter in the data from these systems. The MoS2 coatings showed dramatic decreases in friction coefficients in high-vacuum. Again, this result was not surprising. The unexpected results was in the polymeric systems which displayed decr ease in friction response when humidity was removed from the system. Given the chemical inertness of these systems, it was hard to determine what caused this change. However, more recent work in this field has indicated there is a sensitivity to water for PEEK composites.(54) Pump and Purge Test Results The other test outlined in th e experimental procedures section was the pump and purge experiment. The importance of this experiment was to determine if repeated cycling of the

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87 environments had any effect on the friction respon se of the materials. To evaluate this, two different samples were chosen, the MoS2 with titanium coating and a bulk polymer composite of PTFE with PEEK and MoS2. Both of these samples were run in alternating environments to determine the effects of cycling the environment. The figure below illustrates the results for each sample. Figure B-1 Pump and purge data results for a thin coating and bulk polymer. (a) MoS2 with titanium and (b) PTFE with PEEK and MoS2 In both cases, the material began at high v acuum and was exposed to humid air then the cycle was repeated. The samples displa yed a similar response, although the MoS2 with titanium was much more dramatic, both saw the lowest fric tion under high vacuum in the initial cycles. The friction immediately increased upon exposure to humid air, and after pumping the system back down, the friction did not recover to its initial value, indicating some oxidation or permanent deformation of the system occurred wh ile running in a humid environment. In the polymer experiment, it appears the system did not reach an equilibrium state prior to swapping the environment. The friction was clearly affected by the humidity in the system, but it is not possible to determine the extent of the change as the friction seems to be increasing continually during the experiment. One possi ble explanation for this is a continual deformation of the surface throughout the test due to the high contact pressures. This results in a more conformal contact, a larger real ar ea of contact and ultimate ly higher friction forces.

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88 APPENDIX C AUGER ELECTRON SPECTROSCOPY OF COATINGS Of the coatings tested, two were chosen to have Auger electron spectroscopy performed on them. These were the MoS2 with titanium coating, mainly because it is designed to be environmentally insensitive, and the MoS2 with Sb2O3 and graphite, because this coating is designed to rely on different constituents to pr ovide solid lubrication in differing environments. The samples used were from the 10,000 cycle single environment experiments. Two tracks were used for each sample; the first was run in humid air for 10,000 cycles and the second in dry nitrogen for 10,000 cycles. In each case, a line scan was performed across the track evaluating different elements. The MoS2 with titanium samples were checked for carbon, molybdenum, oxygen, sulphur and titanium. The MoS2 with Sb2O3 and graphite was tested for carbon, molybdenum, oxygen, sulphur and antimony. In all four systems there were distinctly different signatures. The data shown in all the plots are th e normalized intensities for each element. This was chosen because some elements are more likel y to emit Auger electrons than others skewing the raw data. Sulphur, for example, is one of the elements most likely to emit an Auger electron, so its intensity dominates most of the scans. By normalizing the data, it is possible to look at the intensities of all the atoms inside and outside the tracks more easily. Figure C-1 Relative intensities for several elements plotted vs. position across the wear track for the MoS2 with titanium coating run in humid ai r environment. Auger data indicates a high content of titanium and oxygen (likel y titania) on the surface inside the wear track. The signature of the other elements was relatively constant inside and outside the track.

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89 Figure C-2 Relative intensities for several elements plotted vs. position across the wear track for the MoS2 with titanium coating run in dry ni trogen. Like the humid air data, dry nitrogen data indicates a higher content of titanium and oxygen (lik ely titania) on the surface inside the wear track. However, ther e are also distinct changes in the sulphur and molybdenum signatures inside the track indicating more MoS2 is on the surface. The MoS2 with titanium coating had a distinct signature for titanium and oxygen inside the wear track in both humid air and dry nitrogen. This indicates titania is being formed in the system as the track is worn in. The signals from molybdenum and sulphur remained constant across the track indicating there was still a reasonable amount of MoS2 remaining in the track. The track run in dry nitrogen did have an incr ease in titanium and oxygen inside the track, but there was also an increase in sulphur. While the environment is dry nitrogen, there are always impurities present. These results indicate any oxygen in the environment seems to be trapped by the titanium in the coating, but the pres ence of titania does not impede the MoS2 and may help to prevent its oxidation. Figure C-3 Relative intensities for several elements plotted vs. position across the wear track for the AFRL chameleon coating run in humid air. There is no signal from the molybdenum in this coating. Oxygen and carbon are primarily found inside the wear track while sulphur and antimony are depleted inside the track.

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90 Figure C-4 Relative intensities for several elements plotted vs. position across the wear track for the AFRL chameleon coating run in dry nitrogen. The molybdenum, sulphur and carbon intensities increase inside the wear track. The difference between this scan and the open air tests indicate MoS2 is, in fact, drawn to the surface of the coating during wear. The MoS2 with Sb2O3 and graphite results indicated that the coating does have a stronger signal from MoS2 on the surface running in dry nitrogen as opposed to humid air. The humid air results indicated mostly carbon and oxygen in the wear track. This result is expected based on the hypothesis that the preferred solid lubricant is drawn to the surface. The lack of a signal from molybdenum and the sharp decrease in sulp hur signal inside the we ar track also confirm that carbon (could be graphitic), not MoS2, is the primary constituent on the surface of the material after running in a humid environment.

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91 APPENDIX D GIBBS FREE ENERGY CALCULATIONS The oxidation of MoS2 to MoO3 is a process of importance in this study. Determining which species are primarily involve d in the reaction is key to pred icting when it will occur. To that end, the two species present in humid air that are known to reac t with a variety of materials, Oxygen and water were evaluated as causes for oxidation. A chemical reaction involving water and MoS2 is listed below. This equation is generated solely by balancing the reactants with the products. 2232232 M oSHOMoOHSH (D-1) For this chemical reaction to occur spontaneous ly the Gibbs free energy of the system must be less than 0 kJ/mol. The reaction is assumed to occur at room temperature 298K. The values for enthalpy and entropy for each spec ies is listed in the table below. Table D-1. List of values for enthalpy and entropy for species in (D-1).(55) To calculate the free energy of the equation, the value can be found using equation D-2. HSTG (D-2) The value for the reaction listed in D-1 is calculated as follows: 32222 3222223 23 1000MoOHSHMoSHO MoOHSHMoSHOHHHHH T SSSSSG (D-3)

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92 745420276726 298 7841213163563 1000 G (D-4) 205kJ G mol (D-5) This result indicates the reaction will not occur spontaneously until the temperature is far above the decomposition temperature of the material. Another possible reactio n is that Oxygen present in the environment is causing the oxidation of the MoS2. This reaction is defined in D-6. 22322724 M oSOMoOSO (D-6) Using D-2, and the values from Table D-1, the free energy of this reaction is calculated as follows: 3222 32222427 2427 1000MoOSO MoSO MoOSO MoSOHHHH T SSSSG (D-7) 149011885520 1569921261435 1000T G (D-8) 2249kJ G mol (D-9) Because the free energy of the r eaction is far below zero at room temperature, the reaction is expected to occur spontaneously. A third reaction includes the combination of wa ter and oxygen in the system to oxidize the MoS2. 22232229424 M oSOHOMoOHSO (D-10) Using the same method from D-7 to D-9, the free energy of this reacti on is calculated to be 4177kJ G mol (D-11)

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93 Both water and oxygen can be found in the system in all the experiments listed in this study. These results only strengthen the conclusion that further testing must be performed in vacuum where the partial pressures of every species can be accurately measured.

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98 BIOGRAPHICAL SKETCH Matthew Adam Hamilton was born January 21, 1978 in Orlando, Florida. He spent a majority of his youth in Braden ton, Florida where he attended Southeast High School. In August of 1996, he entered the University of Florida. Subsequently, the football team won its first national championship. A turning point in his a cademic career occurred in the year 2000, when he began doing research with Dr. W. Gregory Sawyer in the Tribol ogy Laboratory. After earning his bachelors degree in May 2001, Matthew continued with graduate studies in the field of tribology. Upon completing his masters degree in May 2003, he took a job with a software company in New York City. After one year, the opportunity to come back to school and complete his doctorate was presented to him by Greg Sawyer. In June 2004, Matthew was readmitted to the University of Florida. He completed his doctorate in June 2007 accompanied by another national championship for the football team and back-to-back national championships for the basketball team. With aspirations of becoming a professor, Matthew accepted a postdoctoral appointment with Dr. Robert Carpick at the University of Pennsylvania to begin in August 2007.