An Investigation into the Tribological Properties of Thermally-Oxidized 6Al-4V
Titanium Coupled with Select Engineering Polymers
by
Matthew E. Lessard
An Engineering Project Submitted to the Graduate
Faculty of Rensselaer Polytechnic Institute
in Partial Fulfillment of the
Requirements for the degree of
MASTER OF ENGINEERING IN MECHANICAL ENGINEERING
Approved:
_________________________________________
Dr. Sudhangshu Bose, Thesis Adviser
Rensselaer Polytechnic Institute
Hartford, Connecticut
April, 2011
© Copyright 2011
by
Matthew E. Lessard
All Rights Reserved
2
CONTENTS
LIST OF TABLES ............................................................................................................. 4
LIST OF FIGURES ........................................................................................................... 5
ACKNOWLEDGMENT ................................................................................................... 6
ABSTRACT ...................................................................................................................... 7
1. INTRODUCTION ....................................................................................................... 8
1.1
Polymer Composite Bearing Materials ............................................................ 10
1.2
The Titanium-Polymer Tribosystem ................................................................ 12
2. THEORETICAL BACKGROUND........................................................................... 16
2.1
The Hydrogen-Assisted Wear Mechanism ...................................................... 16
2.2
Thermal Oxidation (TO) Treatment of Titanium ............................................. 21
3. METHODOLOGY / APPROACH ............................................................................ 26
3.1
Thermal Oxidation (TO) Surface Engineering ................................................ 27
3.2
Polymeric Test Materials ................................................................................. 31
3.3
Wear Testing (Pin-on-Disk Method) ............................................................... 33
4. PIN ON DISK TRIBOMETER ................................................................................. 36
4.1
Design and Configuration ................................................................................ 37
4.2
Data Acquisition System (DAQ) ..................................................................... 44
5. EXPERIMENTAL RESULTS .................................................................................. 47
5.1
Test Disks (Thermally-Oxidized Ti64) ............................................................ 47
5.2
Alpha Case Analysis ........................................................................................ 51
5.3
Polymer Test Pins ............................................................................................ 54
5.4
Wear Testing Results ....................................................................................... 55
6. DISCUSSION AND CONCLUSIONS ..................................................................... 72
7. REFERENCES .......................................................................................................... 75
3
LIST OF FIGURES
Figure 1. SEM Image of Wear Scar on the Test Surface of a 6Al-4V Titanium Test Disk
(Subjected to Dynamic Pin-on-Disk Wear Testing) [7] ............................................ 13
Figure 2. SEM image of Beach Marks on a Worn Ti64 Disk Surface [2] ...................... 14
Figure 3. SEM Image Showing Severe Wear of a Ti64 Test Surface by Third-Body Wear
Mechanism [25] ......................................................................................................... 15
Figure 4. Illustration of Proposed Hydrogen-Assisted Wear Mechanism for the
Tribosystem UHMW/Ti6Al4V [8] ............................................................................ 18
Figure 5. Chart Comparing Wear Rates (as a function of Incubation Time) for Various
Metal-Polymer Tribosystems [7] ............................................................................... 19
Figure 6. The Unit Cell of Rutile TiO2 (Titanium atoms are grey and oxygen atoms are
red). [29]
22
Figure 7. Cross-Sectional Micrograph (a) Showing Alpha Case Layer Detail and Plot
(b) showing Knoop Microhardness for a Thermally-Oxidized Ti64 Sample [27] .... 23
Figure 8. Alpha Case Depth as a Function of Heat-Treatment Temperature for
Commercially-Pure (CP) Titanium, Held at Temperature for 1-2 Hrs [4] ............... 24
Figure 9. Microhardness-Depth Profile for Two 6Al-4V Titanium Articles Subjected to
Heat Treatment in an Air Atmosphere at 850°C for Different Lengths of Time [10]25
Figure 10. A Comparison Showing the Variation in Case Thickness for Furnace-Cooled
samples and Air-Cooled samples as a Function of Treatment Time (A) and
Temperature (B). Figures (C) and (D) Show SEM Image of Oxide Layer CrossSection for Air-Cooled Sample and Furnace-Cooled Sample (respectively) [20] .... 29
Figure 11. The Basic Chemical Structure of the Acetal Copolymer Chain [16] ............. 32
Figure 12. The General Chemical Structure of Epoxy [17] ............................................. 33
Figure 13. General Configuration of the Pin on Disk Wear Test .................................... 33
Figure 14. Pin on Disk Tribometer .................................................................................. 36
Figure 15. Pin on Disk Tribometer Cross-Section View ................................................. 38
Figure 16. Pin on Disk Load Arm Component ................................................................ 40
Figure 17. Load Arm Structural Analysis at Max Anticipated Operating Load
Conditions (Spring FS=180 lbf, FD=24 lbf) ............................................................... 42
Figure 18. Test Rig Proximity (Speed) Sensor ................................................................ 43
4
LIST OF TABLES
5
ACKNOWLEDGMENT
I would like to thank professors Sudhangshu Bose and Ernesto Gutierrez-Miravete for
their guidance and support while developing and compiling this work. I would also like
to thank my mentors and colleagues at Kamatics Corporation for their direction and
assistance, especially Mat Mormino, Mark Broding and Tom Rutledge. I owe many
thanks to my parents who have always supported and encouraged me during trying
times. And above all, I would like to thank my wife Tiffany for her love, support and
patience while I was earning my degree.
6
ABSTRACT
This project examines the tribological properties of the titanium-polymer
tribosystem and evaluates the potential improvements that can be achieved by employing
titanium Thermal Oxidation (TO) surface engineering techniques. The specific wear
rates realized when coupling titanium alloys with various engineered plastics
(polyethylene, thermosets, etc.) in dynamic operation are often found to be significantly
higher than those experienced between the same polymeric materials and other metallic
materials possessing comparable interface characteristics (finish, hardness, etc.). It is
believed that these relatively poor wear characteristics are directly linked to a hydrogenembrittlement phenomenon associated with the titanium substrate; one which results
from the breakdown of certain reactive groups found in many engineering polymers. By
subjecting titanium alloys to specific thermal process treatments in an oxygen-containing
atmosphere, it is possible to develop a dense oxygen-diffused case on the surface of the
titanium alloy. Based on historical testing performed by other researchers, it is believed
that this oxygen diffused layer should be effective in preventing the inward diffusion of
hydrogen polymer byproducts into the titanium bulk, thereby reducing the specific wear
rates significantly in these systems.
The ultimate goal of this project is to evaluate the tribological performance of
6Al-4V titanium as related to three specific engineering polymers, and also the
improvements that can be achieved by applying different types of thermal oxidation
treatments to the titanium interface surface.
7
1. INTRODUCTION
As a structural material, Titanium and its alloys have proven their utility in a variety
of industries throughout modern design; none more so than in the aerospace industry.
The relatively light weight of titanium, coupled with its exceptional mechanical
properties (high strength) and inherent corrosion-resistance make it invaluable as a
structural material in an industry where weight-savings comes at a premium. Despite its
mechanical merits however, the relatively poor tribological properties (wear and friction
characteristics) of titanium tend to limit its application in systems with the potential for
dynamic contact. This is certainly a disadvantage when employing titanium alloys in
kinematic linkages and dynamic load bearing applications, to the point that significant
research and design efforts have been dedicated in recent years to developing coatings
and surface engineering processes to mitigate wear and reduce friction in systems
containing these materials.
Historically, the majority of these titanium surface engineering efforts have focused
on the use of hard coatings and similar technologies to establish a wear-resistant barrier
between titanium components. Hard coatings rely on the deposition of certain materials
(i.e. chrome oxide, tungsten carbide, etc.) on interface surfaces using various application
techniques (thermal spraying, chemical vapor deposition, physical vapor deposition) that
yield a very hard and abrasion-resistant superficial coating. The use of hard interface
coatings is an effective practice in reducing titanium component wear, however, most
are still limited to relatively low bearing pressures in the absence of boundary
lubrication. When dynamic loads increase beyond threshold contact pressures, friction
values often rise to unacceptable levels, leading to excessive wear and premature failure.
This need for titanium interface coatings and materials capable of dynamic
operation under higher contact-pressures often leads designers to consider the use of
polymeric bearing materials (engineered plastics) for those applications where surface
coating technologies fall short. The inherent low-friction characteristics and controlled
wear-rates that can be achieved by coupling polymer composites with controlled mating
surfaces makes them ideal for dry sliding bearing applications that operate at low speeds.
8
Metal-polymer tribosystems (dynamic material couples) are typically designed with
specific materials that yield very low frictional forces and also keep the rate of material
wear (volume loss) very low. These tribosystems are designed such that all material
wear occurs in the polymer material, effectively making the plastic component a
sacrificial element in the assembly and saving the cost of replacing mating metallic
components.
There are a number of metallic and ceramic materials that present a suitable mating
surface for bearing polymers, most of which are characterized by a relatively high
material hardness and fine surface finishes. Historically, it is these characteristics that
have demonstrated the greatest impact on specific wear rates at high contact pressures,
however the composition of the mating substrate also plays an important role in the life
of the bearing system. High-hardness, corrosion-resistant steels are the most common
mating materials for engineered plastics; with a carefully controlled mating surface
finish, very low, uniform polymer wear rates can be achieved in addition to a lowfriction interface. Similar (low) polymer wear rates can be expected when mated with
other materials such as nickel-based and cobalt-based super alloys that maintain similar
levels of hardness and surface finish requirements.
Interestingly however, when mating these same polymer materials against titanium
alloys possessing the same general material properties (similar hardness levels and
surface finish requirements) as the more common stainless steels, the resultant specific
wear rates are often found to be significantly higher for the same operating conditions.
In addition to the increase in specific polymer wear rates, these dynamic titaniumpolymer systems also often result in actual wear (volume loss) of the titanium
countersurface; a condition that is not realized with other mating materials. Given the
significant difference in relative material hardness between titanium alloys and the
mating plastic materials, this wear phenomenon proves quite puzzling.
As discussed later in this paper, it is believed that this unusual wear phenomenon is
a direct result of a hydrogen-embrittlement effect developed in the titanium substrate
9
during operation. It is the breakdown of certain reactive polymer groups found in a
number of engineered plastics that ultimately drives this reaction and the ensuing wear
distress caused in the titanium interface surface. Historical research [7. 8] has shown
that this wear mechanism is prevalent in the 6Al-4V titanium / UHMW polyethylene
dynamic tribosystem and also that the application of certain surface engineering
processes (thermal oxidation) can significantly reduce the wear of titanium surfaces in
this system. It is known that this same wear mechanism also exists in other titaniumpolymer tribosystems and this project will attempt to characterize the wear and friction
properties of three such systems. Additionally, this project also attempts to evaluate the
improvements that can be achieved by employing various sur
1.1 Polymer Composite Bearing Materials
The composite bearing polymer designation encompasses a wide range of
engineered plastics that are specifically designed to function in dynamic systems with
demanding operating and environmental conditions.
Polymer bulk matrices are
modified by the addition of strengthening and self-lubricating ‘fillers’ to
in a bulk polymer matrix is an effective method of reducing relative friction and
specific wear in metal-polymer and ceramic-polymer tribosystems.
When polymer
composites are coupled with metallic materials possessing specific surface
characteristics (high superficial hardness, fine surface finish, etc.), are often realized.
The utility of engineering plastics is realized where applications require structural
materials that possess a level of chemical resistance or reduced component weight that
cannot be achieved by metals or ceramics. It is the tribological properties of these
materials, however that makes them invaluable in applications where dynamic contact
exists between machine components.
10
Significant advances have been made in recent years in the field of polymer
tribology, enabling the use of these materials in systems with very demanding operating
conditions. Of particular interest is the usefulness of polymer materials as solid
lubricants in sliding bearing applications where boundary lubrication is not feasible or
cost-effective. In modern tribology, only a few select materials have been identified as
possessing the unique molecular properties necessary to make for an effective bearing
polymer.
……LEAD IN REQD
To make for a useful tribological material, significant efforts have been made to
increase the wear resistance of PTFE, leading to the development of a number of PTFE
composite materials. These composites can typically be classified according to two basic
categories which describe the general method of modification [XX]; Bulk-modified
composites and Interface-modified composites. Bulk-modified composites employ
‘hard’ filler materials such as ceramics, metals or synthetic fibers directly in the bulk
self-lubricating material (i.e. Polytetrafluoroethylene or PTFE). The function of these
‘hard’ filler materials is essentially to strengthen the polymer matrix, increasing its loadcarrying capacity and wear-resistance. Some of the more common filler materials used to
increase bulk material strength are glass, aramid and carbon (all usually in short fiber
form).
On the other end of the composites spectrum, interface-modified composites employ
a softer self-lubricating material (i.e. PTFE, graphite, molybdenum disulfide) as a ‘filler’
in a harder polymer matrix. The PTFE fillers (also commonly in short fiber form)
provide the composite with the desired low friction properties, while the strength and
toughness of the bulk matrix (some common polymers are epoxy, phenolic and PEEK),
is augmented by additional fillers. Alumina (aluminum oxide) and titanium oxide are
frequently employed as strengthening fillers in interface-modified polymer composites
and their particle size is central to the level of wear resistance that can be achieved using
these fillers. Recent studies [2] have shown that using fillers with particle size on the
nano-scale (termed nanocomposites) can effectively reduce wear rates by one and in
11
some cases, two orders of magnitude when compared to micro-sized filler particles. This
is also true of bulk-modified composites.
The application of engineered plastics and self-lubricating polymers as a barrier
against wear distress (and ensuing fretting fatigue failures) in structural components has
proven an effective technology in the aerospace industry; most notably in bearing
applications. By interfacing ferrous materials possessing specific surface characteristics
(high material hardness / fine surface finishes) with engineered polymer composites, it is
possible to design assemblies that are capable of operating at very high bearing contact
pressures while maintaining very low, controlled wear rates and friction coefficients.
The high strength-weight characteristics of titanium alloys would therefore make them
an ideal candidate for such applications, however the wear characteristics of this material
are often prohibitive.
1.2 The Titanium-Polymer Tribosystem
In most metal-polymer tribosystems, the relationship between material wear and
hardness agrees with the standard model for wear of materials that was proposed by John
F. Archard [18] in the early 1950’s (often referred to as the Archard Equation).
According to Archard, for most isotropic materials, the rate of wear in a material is
found to be inversely proportional to the hardness of that material.
Where:
v = Material volume loss
k = Wear coefficient
W = Normal load
H = Material Hardness
That is to say in a simplified model that, as the hardness of the material is increased,
the specific wear rate of the material will decrease. This is typically also true of the
wear rates of the mating polymer materials in the same tribosystems. As the hardness of
the metallic material in this couple is increased, the specific wear rate of the polymer
material is most often also found to decrease. This relationship has been effectively
12
comprehensively verified by W. Wieleba [19] through his work using PTFE composites
and various steel mating surfaces.
The titanium-polymer tribosystem however, appears to be an exception of sorts to
the standard wear model that was proposed by Archard. Numerous studies have been
conducted [2, 3, 7, 8] investigating the tribological behavior of titanium coupled with
several different engineering polymers using varying operating conditions and
researchers consistently witness the same type of wear phenomenon regardless of
operating parameters (load, speed, etc.). In most every case, the specific wear rates
developed in the polymer materials were much higher than expected and more notably,
the (much harder) titanium countersurfaces also displayed significant wear. In general,
the increase in the wear rate of the plastic material could possibly be explained as strictly
being a function of the dynamics between the titanium microstructure and that of the
polymeric material, however the wear mechanism developed in the titanium substrate is
much more difficult to explain.
Figure 1. SEM Image of Wear Scar on the Test Surface of a 6Al-4V Titanium Test
Disk (Subjected to Dynamic Pin-on-Disk Wear Testing) [7]
The relative hardness (and correspondingly, the strength) of alloys such as 6Al-4V
(the most common of the titanium alloys) is two and in some cases, three orders of
magnitude higher than the engineered plastics with which they are mated.
When
compared against ultra high molecular weight polyethylene (UHMW/Ti64 is a
tribosystem that is commonly found in the medical industry), the relative hardness of
6Al-4V titanium is nearly 700 times greater [8]. Given this large disparity in material
13
hardness, one would tend to expect that the softer polymeric material would exhibit all
of the component wear in this tribosystem, however experimental testing has
demonstrated that during dynamic operation under load, the titanium component
consistently displays significant material wear.
For instance, the rate of titanium
material wear generated during dynamic operation against UHMW polyethylene plastic
is quite often comparable to the titanium wear rates that develop when running against
significantly harder tool steels and ceramic materials using the same operating
conditions [3].
An examination of the actual wear scars that develop on titanium surfaces during
testing against these engineered plastics suggest that there are some abnormal wear
mechanisms at work behind this strange wear phenomenon; characteristics that are not
typically found in steel-polymer tribosystems. This is evidenced by the presence of
certain wear pattern characteristics that indicate a type of micro-fretting fatigue failure
mechanism.
One such characteristic that is frequently discovered post-test is the
presence of ‘beach marks’ or superficial micro-cracks oriented transversely across the
wear zone (perpendicular to the sliding test direction). This micro-cracking pattern
indicates a lack of material toughness at the surface of the material which translates to
micro-crack propagation and liberated wear particles.
Figure 2. SEM image of Beach Marks on a Worn Ti64 Disk Surface [2]
14
Another titanium surface failure mechanism that often manifests itself during
testing at elevated contact pressures is the presence of ‘wear craters’ or localized
liberation of larger titanium wear particles in a scarred wear field.
This wear
characteristic is sometimes difficult to detect however as the liberated particles often
become trapped in the wear zone and a type of third-body wear mechanism is generated.
Liberated wear particles trapped in the wear path will sometimes become embedded in
the softer mating counterface, increasing the depth of axial wear scars during additional
testing through a plowing action.
Figure 3. SEM Image Showing Severe Wear of a Ti64 Test Surface by Third-Body
Wear Mechanism [25]
These are just two distinct characteristics that are frequently identified in the
wear scar of a titanium surface following dynamic testing against certain polymer
countersurfaces. There are a number of other unique wear characteristics that also
develop in these tribosystems, many of which are influenced by operating conditions
such as speed, contact pressure, temperature, etc. An experienced metallurgist could
certainly theorize on the root cause for each individual wear characteristic, however, it is
not the goal of this project to develop a taxonomy of titanium wear mechanisms. Rather,
this work hopes to identify a common link between these failure modes and to test a
surface engineering process that has the potential to significantly reduce wear in these
systems.
15
2. THEORETICAL BACKGROUND
2.1 The Hydrogen-Assisted Wear Mechanism
One theory to explain the strange wear phenomenon that exists in titaniumpolymer tribosystems was proposed by three researchers in 2001 during an experimental
investigation of the wear performance of 6Al-4V titanium coupled with UHMW
polyethylene [8]. Their work originally initiated in 1999 [7] as a study of various
titanium surface engineering processes that might be used to reduce the wear rate of
UHMW polyethylene when operated against titanium materials.
This is a common
material combination in medical total joint replacement prosthesis applications and
premature wear of the plastic component in these systems is a constant concern. Using
standard pin-on-disk wear testing methods, H. Dong and his associates were able to
successfully demonstrate significant gains in the friction and wear properties of UHMW
coupled with Ti64 by introducing different coatings and surface engineering techniques;
specifically, diamond-like coatings (DLC), thermal-oxidation (TO) and oxygen diffusion
(OD) surface treatments.
During the course of their testing however, these researchers encountered an
unexpected wear phenomenon which they initially found difficult to explain.
To
establish a control baseline for comparison against each of the individual samples, wear
testing was first performed using an un-treated 6Al-4V titanium test disk (dynamically
tested against a UHMW polyethylene test pin). After completing this baseline test, they
discovered that the titanium test surface had sustained significant wear distress in the
form of deep, wide grooving (scarring) while the much softer plastic test pin maintained
a relatively low wear rate. During subsequent testing with the modified test samples
(DLC, TO and OD), none of the surface-modified titanium components exhibited any
wear.
Initially, these researchers attempted to explain this wear phenomenon as
characterized by a three-body abrasive wear mechanism, perpetuated by a poor ionic
16
attraction between the titanium substrate and the thin oxide layer that forms naturally on
titanium surfaces that are exposed to air. They suggested that this natural surface oxide
layer (very thin, typically 5-10 nm thick) became liberated during operation against the
UHMW test pin and particles were subsequently trapped at the test interface in the
polymer material, generating the abrasive wear mechanism. Following an additional two
years of wear testing with Ti64-UHMW material couples, this team found new evidence
to support a different type of theory for this wear condition; hydrogen-assisted wear.
The hydrogen-assisted wear mechanism appears to originally have been proposed
by a team of Russian researchers in the late 1970’s [13], and then with the first English
publication in 1993 by A. L. Zaitsev [11] while examining the tribological performance
of several polymeric materials coupled with tungsten carbide-cobalt coatings. During
testing, Zaitsev noticed a trend wherein hard alloys (primarily tungsten and cobalt-based
materials) experienced higher than expected wear rates when coupled in dynamic contact
with polymers such as polycaproamide (PCA), phenol formaldehyde polymer (PF) and
EDC epoxy compound. All of these materials are found to contain reactive chemical
groups such as amide, ether and hydroxyl groups in their polymer structure.
The theory of hydrogen-assisted wear that was proposed by Zaitsev and then
advanced considerably by Dong [8] attempts to explain this abnormal metal-polymer
wear relationship as being a function of the reactive chain groups found in these
polymers. During the dry sliding action between these materials and their metallic
counterparts under load, a significant amount of dynamic energy is dissipated in the
form of heat. Dong and Zaitsev suggested that this intense localized heat, coupled with
the plastic deformation and strains that occur at the test interface are enough to cause a
breakdown of the hydrogen-containing amide and hydroxyl groups found in these
engineered plastics. During continued dynamic operation, the high heat retained at the
test interface aids the inward diffusion of this excess hydrogen into the (titanium) bulk
material substrate.
17
Titanium is a material that is known to be susceptible to hydrogen-embrittlement,
a phenomenon that reduces the toughness (general resistance to cracking) and ductility
of a material. The natural propensity of hydrogen atoms for titanium atoms is also a
driving function in this relationship; the hydrogen atoms tend to work their way into the
titanium bulk and collect at grain boundaries. As these hydrogen atoms accumulate at
these locations and subsequently develop into titanium hydrides, their growth causes
subsurface pitting and microcracks in the titanium bulk material. Subsequently, the
interfacial shear forces generated by the sliding action of the plastic test material across
the outer surface of the titanium test article subsurface fracture of the titanium material
(liberation of wear debris particles). Figure 4 provides a graphical illustration of the
hydrogen-assisted wear mechanism as proposed by these researchers.
Hydride Growth and
Surface Cracking
Hydrogen Diffusion
Third Body (Abrasive) Wear
Mechanism - Wear Particles
Embedded in Mating Surface
Hydride Formation
Figure 4. Illustration of Proposed Hydrogen-Assisted Wear Mechanism for the
Tribosystem UHMW/Ti6Al4V [8]
In an effort to validate this theory, Dr. Hanshan Dong and his associates [8]
conducted a series of dynamic pin-on-disk wear tests using the Ti64/UHMW material
couple, and set to work characterizing the associated wear mechanisms by analyzing the
morphology of both the titanium test surfaces and UHMW plastic countersurfaces.
Testing was conducted using standard unmodified 6Al-4V mill product (hot-rolled, mill
18
annealed bar) to establish the base wear relationship with UHMW and testing was also
conducted with Ti64 samples that had been subjected to a unique thermal oxidation
process. Additionally, annealed 316L stainless steel samples were subjected to the same
wear test so that a comparison could be made between the Titanium/UHMW couple to
that with a stainless steel product of lower material hardness. All testing was conducted
using water-lubricated operating conditions (simulating periprosthetic applications) and
Figure 5 below provides a summary of the time that was required to generate a certain
level of wear in each sample (termed Incubation Time; a longer incubation time indicates
a lower material wear rate). Correlating incubation time with specific wear (of the
metallic mating surfaces), it can be seen that the wear rate of the as-polished titanium
test surface was significantly higher than the 316L test surface (approx 15x-20x higher)
even though the hardness of the stainless material was about 2/3 that of the titanium test
material
Figure 5. Chart Comparing Wear Rates (as a function of Incubation Time) for
Various Metal-Polymer Tribosystems [7]
These results seem to agree well with the results obtained by other researchers
performing similar wear tests with titanium alloys; excessive wear damage sustained by
the titanium test surface during dynamic operation against a much softer engineered
plastic material. After completing these tests, Dong (et al.) performed an extensive
analysis for both the un-modified and thermally-oxidized samples using transmission
19
electron microscopy (TEM) to inspect the test surface morphology. Glow discharge
spectroscopy (GDS) was also used to analyze the depth profile of elements in the wear
track post-test.
During the course of their analysis, these researchers discovered that a significant
concentration of titanium hydrides (TiHX) were present in the outer titanium oxide layer
and more importantly, in the α-phase titanium bulk – hydrides that did not exist prior to
testing. The formation of titanium hydrides below the direct interface layer indicates
that the diffusion of hydrogen products into the titanium bulk could well have
contributed to the deep abrasive wear that was experienced during these tests. This
condition was detected in the un-modified titanium test samples as well as the thermallyoxidized and oxygen-diffused test samples. However, the percent composition of TiHX
found in the α-Ti bulk of the thermally-oxidized samples was significantly lower than
the percentage that was detected in the un-modified titanium samples. This stands to
reason as testing has demonstrated that the naturally-occurring outer titanium surface
oxide layer serves as a fairly effective barrier against hydrogen diffusion (reduces the
diffusivity of hydrogen into titanium) [12].
It was based on these findings that Dr. Hanshan Dong (et al.) proposed the theory of
hydrogen-assisted wear as potentially being the driving function behind this abnormal
titanium-polymer wear relationship. The presence of subsurface titanium hydrides in
these tested articles appears to be a strong indicator that hydrogen has a significant
influence on the wear of titanium alloys by softer plastic materials. This theory was
established using wear results generated with the 6Al-4V titanium–UHMW polyethylene
tribosystem, however it also pertains to tribosystems involving titanium coupled with
other engineered plastics containing certain polymer chain groups.
20
2.2 Thermal Oxidation (TO) Treatment of Titanium
Inherently, titanium materials have excellent corrosion-resistant properties and are
not easily compromised by attack from acidic media. These protective properties are
afforded by the presence of a thin oxide layer that forms naturally on the surface of
titanium materials over time as they are exposed to an oxygen-containing atmosphere
(air). This oxide layer is typically very stable and adherent, but also is commonly very
thin. After long-time exposure of newly-machined titanium to a typical air atmosphere,
this outer oxide layer will only grow to be about 3-4 nm (.1 μ-in.) thick [7]. Many
researchers believe that the poor wear relationships exhibited by titanium tribosystems is
primarily attributed to the removal of this thin oxide layer by the dynamic interaction
(relative sliding motion) with a mating surface and also, that a more robust boundary
layer applied to this interface could help to improve this condition.
Dr. Hanshan Dong demonstrated in his research efforts [7, 8] that the application
of thermal oxidation treatments to titanium-base alloys can significantly improve the
tribological properties of these materials when dynamically coupled with UHMW
polyethylene. The wear phenomenon that affects this material couple however, is also
known to occur in other titanium-polymer tribosystems [5]; systems involving
engineered plastics containing similar reactive polymer end groups. Based on the theory
of Hydrogen-Assisted wear, the application of titanium TO treatments should also be
effective in mitigating wear in these dynamic systems. (…and the development of
different titanium TO treatments is the focus of this project…)
The thermal oxidation process is performed much like the name implies, titanium
articles are subjected to a heat treatment process in an oxidizing atmosphere. The
titanium responds to this environment by developing an oxygen-rich, case hardened
outer surface layer which is often referred to as alpha case. This alpha case layer
generally consists of two modified regions or layers:
21
- Oxide Surface Layer - The outermost layer is
typically very thin (usually on the order of a few
micrometers in thickness) and consists primarily
of TiO2 (titanium dioxide).
This usually
manifests as Rutile, the most compact and stable
of the TiO2 polymorphs (the other two being
Anatase and Brookite) and its presence on the
surface of the titanium article significantly
improves corrosion resistance of the titanium
article. The relative hardness of this outermost
Figure 6. The Unit Cell of
Rutile TiO2 (Titanium atoms are grey
and oxygen atoms are red). [29]
layer is also much higher than that of the titanium
bulk (TiO2 being a ceramic in its basic form).
- Oxygen-Diffused Zone – During the heat treatment process, oxygen also tends to
diffuse downward into the bulk titanium, forming an oxygen-diffused (TiO) region
directly below the outer oxide layer. This TiO region is comprised primarily of αphase (hexagonal close-packed) titanium microstructure due to the alphastabilizing nature of oxygen in a titanium matrix (hence the designation alphacase). The solid-solution hardening effect of this oxygen diffusion process is also
considerable. A microhardness check performed on a thermally-oxidized titanium
samples will often show an increase in relative hardness of around 2-3 times that of
the titanium bulk material. Depending on thermal atmosphere conditions, this
oxygen-diffused region can extend anywhere from 2-300+ μm (.00001-.012 in.)
down into the titanium.
22
Figure 7. Cross-Sectional Micrograph (a) Showing Alpha Case Layer Detail and
Plot (b) showing Knoop Microhardness for a Thermally-Oxidized Ti64 Sample
[27]
The physical properties of this alpha case layer, as well as the respective depth of
penetration and microstructural composition are all highly influenced by the thermal and
atmospheric controls employed during thermal oxidation treatments and a wide range of
results can be achieved by varying these parameters. For example, the relationship
between heat-treatment temperature and resultant depth of penetration do not relate
linearly; as heat treatment temperature is increased, an exponential increase in the actual
depth of alpha case is realized. This is a function of the effect that temperature has on
the diffusivity of oxygen into titanium and is also influenced by the beta-transus
temperature of the material.
The beta-transus is the temperature at which the
microstructure of titanium converts from an alpha-phase (hexagonal close-packed)
microstructure to a beta-phase (body-centered cubic) microstructure. The diffusivity of
oxygen into a fully beta material at temperature is higher than in an alpha microstructure.
The beta-transus for commercially pure titanium is around 880 °C (1615 °F) and is
different for the various alloys. Figure 8 shows this relationship between depth of case
and heat-treatment temperature.
The majority of thermal oxidation treatments are conducted at a heat-treatment
temperature below the beta-transus for that specific alloy to reduce the impact on the
general mechanical properties of the titanium. Processing titanium alloys for extended
23
lengths of time above the beta transus can sometimes cause excessive precipitation of
alloying elements out of the solid solution, reducing the strength and toughness of the
material. Additionally, crossing the beta transus temperature initiates the transformation
process of the titanium matrix from alpha to beta phase and the transition back to alpha
phase (cooling) has the potential to result in an undesirable grain structure or other
anomalies such as retained beta / beta fleck (conditions that are known to affect material
toughness).
Figure 8. Alpha Case Depth as a Function of Heat-Treatment Temperature for
Commercially-Pure (CP) Titanium, Held at Temperature for 1-2 Hrs [4]
The hardness-depth profile in the oxygen-diffused zone of the alpha case layer is
primarily attributed the relative density of oxygen atoms in the titanium solid solution
(increasing density correlates to increasing hardness). In addition to the heat-treatment
temperature, the actual time held at temperature, the oxygen content of the furnace
atmosphere and the method of cooling all have an effect on the resulting hardness
gradient in this region. With respect to the dwell time (time held at temperature), Figure
9 shows a hardness-depth profile for two 6Al-4V titanium samples that have been
subjected to a heat treatment process at 850 °C in an air atmosphere for different lengths
of time.
24
Figure 9. Microhardness-Depth Profile for Two 6Al-4V Titanium Articles Subjected
to Heat Treatment in an Air Atmosphere at 850°C for Different Lengths of Time
[10]
Significant work has been accomplished in recent years [XX-XX GROUP] to
increase the understanding of titanium thermal oxidation treatments and the tribological
improvements that can be achieved by employing these surface engineering techniques
in specific operating environments. It is clear from the wide range of results obtained
though experimental testing that the
however, the accurate selection of process parameters will make or break
titanium interface surfaces. A wide range of properties have been…
interface surfaces by carefully controlling the …..XXXXXX
25
3. METHODOLOGY / APPROACH
The theory of ‘hydrogen-assisted’ wear that was proposed by Dr. Hanshan Dong
does much to explain the wear phenomemon behind the titanium-polymer tribosystem
and his work with UHMW [8] offers some very conclusive findings. However, little-tono research has been published involving other titanium-polymer tribosystems [5] that
have also been known to develop these abnormal wear patterns.
Two engineered
polymers that have also been known to cause excessive damage to mating titanium
surfaces during dynamic interaction are acetal copolymer and epoxy compound. The
latter is frequently employed as a bearing polymer in a composite form (combined with
various solution-strengthening and friction-reducing fillers such as PTFE) while acetal
copolymer is often used in its un-modified form as a bearing plastic.
It is the goal of this project to experimentally evaluate the tribological
improvements that can be achieved with the Ti64-acetal and Ti64–epoxy tribosystems
by employing various thermal oxidation techniques. Using an ASTM standard dynamic
wear testing approach, this project will attempt to characterize the wear and friction
properties of these polymer materials in addition to UHMW polyethylene (basis for
comparison against historical research) as operated against titanium samples that have
been subjected to four unique thermal oxidation processes.
To establish a control
baseline for comparison, each of these polymer samples will also be tested against an
unmodified 6Al-4V titanium test sample.
Standard metallurgical characterization techniques are used to determine the
properties of the alpha case product developed during each thermal oxidation process. It
should be noted, however that the equipment required to detect the presence of titanium
hydrides in the tested titanium surface are not available and this project is strictly
26
3.1 Thermal Oxidation (TO) Surface Engineering
The range of case properties and titanium microstructures that can be developed by
varying the control parameters applied in a thermal oxidation treatment is extensive. A
few studies published in recent years [X,X] have attempted to characterize the wear
characteristics of an array of different titanium TO-treated samples, however these
studies were conducted using a 6Al-4V substrate and different steel / ceramic mating
materials. The titanium-polymer tribosystem is inherently different and it is difficult to
say whether the tribological improvements witnessed using different steel and ceramic
countersurfaces would translate over to polymer tribosystems. Each system is unique
and should really be considered on a case-by-case basis.
The ultimate goal of this project is to establish a baseline characteristic wear
relationship for three different titanium-polymer tribosystems and then to evaluate the
improvements in wear and friction that might be achieved by applying titanium thermal
oxidation treatments. Four different thermal oxidation treatments were selected for this
project based on a number of criteria. In each case, consideration was provided to:
-
Case microhardness profile
-
Depth of penetration
-
Composition of the outermost boundary surface layer
-
Adhesion between surface layer and oxide layer
-
Resultant grain structure of the Ti bulk
-
Ratio of boundary surface layer thickness to OD zone
These case properties were all considered while developing the thermal control
parameters for each process while also keeping an eye on the impact that each has from
a manufacturing ‘producibility’ standpoint. For example, parameters such as excessive
time held at temperature could potentially pose a problem if these TO treatments were
applied in a full-scale production environment.
27
Thermal Oxidation Treatment - Case #1
Heat Treat Temperature:
Furnace Atmosphere:
Time at Temperature:
Cooling:
850 °C (1562 °F)
Air
2 Hrs
Air cooled
Case #1 would be considered the most straightforward TO treatment solution from a
manufacturing standpoint; short time at temperature, no furnace atmospheric controls
and air cooling. This treatment has shown favorable results in the thermal oxidation
efforts performed by Dr. Jun Qu and Dr. Peter Blau [21, 23, 27]. Using this process,
they successfully demonstrated that the wear resistance of Ti64 could be increased
significantly in different steel-titanium systems using both lubricated and un-lubricated
operating conditions.
Thermal Oxidation Treatment - Case #2
Heat Treat Temperature:
Furnace Atmosphere:
Time at Temperature:
Cooling:
850 °C (1562 °F)
Air
5 Hrs
Furnace cooled
Case #2 uses the same heat treatment temperature and atmosphere that was used in Case
1 but extends the oxidation time and more importantly, changes the method of cooling.
Research has shown [20] that cooling the thermally-oxidized samples in a furnace as
opposed to an air-cool method can significantly improve the mechanical properties of
the case layer as well as the depth of penetration. The strength of adhesion between the
oxide layer and the substrate is also improved significantly which most often translates
to improved wear resistance.
28
Figure 10. A Comparison Showing the Variation in Case Thickness for FurnaceCooled samples and Air-Cooled samples as a Function of Treatment Time (A) and
Temperature (B). Figures (C) and (D) Show SEM Image of Oxide Layer CrossSection for Air-Cooled Sample and Furnace-Cooled Sample (respectively) [20]
Thermal Oxidation Treatment - Case #3
This is a 2-stage thermal treatment; the parts are initially treated in a standard air
atmosphere for the stated time and then in a vacuum for a given time, followed by
furnace cooling (under vacuum). Number (1) is standard air atmosphere, number (2)
indicates vacuum atmosphere
(1)
(1)
(1)
(2)
(2)
(2)
(2)
Heat Treat Temperature:
Furnace Atmosphere:
Time at Temperature:
Heat Treat Temperature:
Furnace Atmosphere:
Time at Temperature:
Cooling:
850 °C (1562 °F)
Air
25 minutes
850 °C (1562 °F)
Vacuum
20 Hrs
Furnace cool under vacuum
Case #3 represents a very different type of thermal oxidation treatment as compared
against the other ‘traditional’ TO processes evaluated in this project. This thermal cycle
29
employs a unique 2-stage process that is designed to develop a very deep and adherent
oxygen-diffused case layer while minimizing the resultant outer surface TiO2 oxide
layer. The first stage of this process (850 °C / air) initiates the formation of the TiO2
outer surface oxide layer much like the other processes tested in this project. However,
by subsequently subjecting the oxidized samples to continued heat treatment in a
vacuum atmosphere (second stage), a different type of diffusion mechanism is initiated
which forces the oxygen deeper into the substrate. The low partial-pressure of oxygen in
the vacuum atmosphere most likely causes the following reactions:
Ti(s) + O2(g) = TiO2(s) (1)
TiO2(s) + Ti(s) = 2O (dissolved in Ti) (2)
which helps to explain the dissolution of the surface oxide layer (that was generated
during the first stage) and the resultant increase in absorption of oxygen into the
titanium.
Using this unique thermal treatment process, researchers have successfully
demonstrated [10] that oxygen-diffused case depths of 300+ μm (.012”) can be achieved
(as compared to around 100-120 μm using ‘traditional’ TO treatments). Based on the
theory of hydrogen-assisted wear in titanium-polymer tribosystems, it is believed that
the depth of the oxygen-diffused layer could directly be related to the wear resistance of
the sample.
Thermal Oxidation Treatment - Case #4
Heat Treat Temperature:
Furnace Atmosphere:
Time at Temperature:
Cooling:
650 °C (1202 °F)
Air
48 Hrs
Furnace cooled
The goal of thermal oxidation case #4 was to focus more on the tribological properties of
the outer TiO2 oxide scale, while placing less emphasis on the depth and hardness of the
TiO oxygen-diffused zone. This treatment takes place at a much lower temperature (650
°C) than the other three treatments and accordingly, requires a much longer time at
temperature to develop the desired case properties. Experimental research performed by
30
S. Kumar et al. [30] using 6Al-4V titanium suggests that a dense TiO2 layer of
exceptional superficial hardness (higher than reported for most other TO treatments) is
possible.
Kumar’s research however, focused on the microstructural and
electrochemical properties of various TO treatments and less on tribological properties
(no wear performance data reported).
3.2 Polymeric Test Materials
As noted above, the focus of this project is on the tribological properties of surfacemodified titanium coupled with three different engineered plastic materials commonly
used in tribological applications:
-
Ultra High Molecular Weight (UHMW) Polyethylene
-
Acetal Copolymer Resin
-
Epoxy Compound
In their engineered forms, each of these polymer materials can be successfully
employed as a low-friction barrier against wear between metallic components, however
each is also known to exhibit a strange wear relationship with titanium alloys (discussed
in Section 1.2). It has been proposed [8] that this titanium wear phenomenon (hard
alloys damaged by operation against soft polymers) is directly related to a hydrogenembrittlement effect in the titanium that is caused by the breakdown of certain reactive
groups found in various polymer materials; the so-called hydrogen-assisted wear
mechanism.
To date, the only tribosystem to be tested and analyzed for metallurgical evidence to
directly support this theory is the UHMW-Ti64 tribosystem. Neither acetal copolymer
or epoxy compound have been subjected to the regimen of wear tests and subsequent
analysis that UHMW has seen, however both of these materials contain polymer chain
groups that were identified by researcher H. Dong [8] and A. L. Zaitsev [11] as
potentially being drivers in the hydrogen-assisted wear mechanism. Dong and Zaitsev
have suggested that the destruction of amide, ether or hydroxyl groups found in the
31
polymer material ultimately results in the reduction of toughness in the titanium
material.
Acetal copolymer is a linear (unbranching) polymer that is formed by the
copolymerization of trioxane with small amounts of comonomer.
The resulting
oxyethylene groups (ether group family) are formed by random carbon-carbon links
throughout the polymer chain (Figure 11). These C-C bonds are the link that provides
this polymer with a relatively stable chain structure.
However, once subjected to
degradative conditions such as high strain, excessive heat (etc.), the chain will
depolymerize (breakdown) or ‘unzip’ – releasing H-C-H alkanes until it hits a C-C bond
[16]. The mechanics of the depolymerization process associated with acetal copolymer chain
therefore appear to meet the requirements proposed in the hydrogen-assisted wear model.
Figure 11. The Basic Chemical Structure of the Acetal Copolymer Chain [16]
Epoxy is a thermosetting resin that is formed by the copolymerization of two
primary components, an epoxide chain (epoxy resin component) and a polyamine
monomer ‘hardener’ component. The polymerization or ‘curing’ process of epoxy is
characterized by extensive polymer cross-linking and the end result is a fairly complex
chain structure. There are quite a few epoxy resins in use today, all with varying chain
structures evolved through different cross-linking systems. All of these compounds
however, appear to have a common trait, hydroxide groups exist in the base polymer
chain, connected by a single carbon-oxygen bond.
32
Figure 12. The General Chemical Structure of Epoxy [17]
The cross-linking action inherent to the curing process of epoxy compound is
effective in impeding the depolymerization of most epoxy chains, however the
hydrogen-containing hydroxide groups are typically released during this process [14,
17]. This again fits with the hydrogen-assisted wear model as proposed by Dong and
Zaitsev with respect to the titanium-epoxy tribosystem.
3.3 Wear Testing (Pin-on-Disk Method)
One of the most widely accepted methods for evaluating the wear and friction
properties of different material couples is the Pin-on-Disk test. Just as the name implies,
this test method consists of a test ‘disk’ component and a test ‘pin’, oriented such that
the axis of the pin is oriented perpendicular to the flat surface of the disk (and offcenter). With a constant load applied to the pin (along its axis), the wear test is then
conducted by rotating the disk about its center (while the pin remains static). The test
continues until the pin has completed a pre-defined number of cycles or ‘distance
traveled’ in relative motion.
Load applied to test pin
FN (Axial direction)
Wear Track
(Volume Loss)
Test Pin
(UHMW Polyethylene)
(Acetal Copolymer)
(Epoxy Compound)
Test Disk
(Unmodified 6Al-4V Titanium)
(Thermally-Oxidized Titanium)
Disk is rotated about its center
causing circular wear path
V(s)
Figure 13. General Configuration of the Pin on Disk Wear Test
33
As the test disk cycles through revolutions and the test pin continues to travel over
the same circular path under load, the test samples (either the pin, the disk or both based
on the materials being tested) will accumulate wear which is reflected in the form of lost
material (volume loss). After the test is completed (total required distance traveled is
achieved), both the test disk and the test pin are measured to quantify the amount of
material that has worn away. In terms of the test disk component, the total wear can be
calculated by physically measuring the width and depth of the resultant wear scar.
Often, a profilometer is required to measure the depth due to the rough, non-uniform
profile at its root. After testing, wear scars are also commonly very shallow and difficult
to record using standard measurement tools (drop gauge, etc.). Lost volume in the test
pin component is more easily calculated using standard measurement tools (calipers,
micrometer).
Alternatively, the lost material volume in each completed test sample can be
verified by first cleaning, and then weighing the samples using precision scales. The
rate of wear in each test component is then typically reported as a function of distance
traveled and effective pressure. This is commonly referred to as the Specific Wear Rate
for that material or the Wear Coefficient. According to Bhushan [1] the wear coefficient
(K) is calculated as:
With wear rates commonly specified in units of (mm3/N·m) or (in3/lbf·ft).
This Pin-on-Disk test method was determined to be the most suitable approach for
this project based on a number of criteria:
Time Required to Test – The standard testing parameters associated with POD
tests allows for quick turnover. Also, the relatively simple configuration of
the test samples means less time required for manufacture and setups are very
quick.
34
Cost – The overall size of the required test samples is small; geometry is simple
and easy to manufacture.
Reliability – The straightforward configuration of this test limits the potential for
error caused by external and system-related variables
Similarity – This also appears to be the method that was selected for each test
case in all of the research papers published on titanium thermal oxidation.
Maintaining similar operating parameters allows for direct comparison of test
results.
It should also be noted that a Pin-on-Disk tribometer (test apparatus) was not
readily available and it was necessary to design and manufacture one prior to testing.
The ability to manufacture a tribometer in a short amount of time that would provide
reliable and repeatable results with data acquisition capabilities also factored heavily in
the test selection process. The design of this test apparatus is discussed in detail in
Section 4.
35
4. PIN ON DISK TRIBOMETER
The wear testing element of this project was central to verifying and validating the
impact that titanium TO treatments have on the tribological properties of titaniumpolymer tribosystems. As noted in Section 3.3, the ASTM standard Pin-on-Disk test
method was identified as the most appropriate approach to characterize relative wear and
friction properties, however a Pin on Disk testing apparatus was not readily available. A
number of commercial Tribometer systems were evaluated, however all were found to
be cost prohibitive (systems such as the Falex ISC cost $40K+).
It was therefore necessary to design and manufacture a cost-effective Pin-on-Disk
testing apparatus from scratch; one that was capable of producing the desired dynamic
interface and also provided the data acquisition capabilities necessary to record dynamic
friction data.
Figure 14 depicts the test rig which was completely designed,
manufactured and tested during the first six weeks of this project.
The following
sections briefly describe the design approach, features and capabilities of this Pin on
Disk testing apparatus that was ultimately used to conduct all of the wear testing
outlined in this paper.
Figure 14. Pin on Disk Tribometer
36
4.1 Design and Configuration
The apparatus was originally designed using the ASTM standard for Pin on Disk
testing (ASTM G99) as a guideline for functional requirements. This specification is
fairly broad concerning the actual configuration of the apparatus, constraining only the
system level requirements (i.e. specific pin-to-disk interface and alignment requirements,
relative motion, etc.). The actual method of test specimen constraint, load application
and power transmission are all left to the discretion of the designer so initial efforts were
focused on a cost-effective solution that would provide consistent and reliable results.
As directly specified in ASTM G99 Section 3.1, the Pin on Disk wear test consists
of two test specimens; a pin with a radiused or flat tip which is positioned perpendicular
to the other specimen, usually a flat circular disk. The test machine causes either the disk
specimen or the pin specimen to revolve about the disk center. In either case, the sliding
path is a circle on the disk surface. The plane of the disk may be oriented either
horizontally or vertically. The pin specimen is pressed against the disk at a specified
load, often by means of an arm or lever. Load is then applied to the pin specimen
(normal to the face of the disk) and maintained throughout the test.
The test apparatus designed for this project is configured with a load arm / disk
spindle-type configuration. The test disk component is fastened directly to a cylindrical
‘spindle’ using a shoulder screw to ensure accurate alignment and concentricity between
the test disk and the spindle. The spindle detail is machined from 17-4 PH stainless steel
(condition H1150) and is supported and constrained in X, Y and –Z directions by two
ABEC1 sealed ball bearings. These bearings provide the supporting reaction force
against thrust loads applied normal to the disk test surface and also allow rotation of the
spindle/disk assembly about the axis of the spindle. Ball bearings were selected over
roller thrust bearings for this application as they introduce less frictional drag force into
the system (the calculated increase in effective drag force is less than 1% of the actual
test force in the drag force – within the limits of the strain gauge measurement error).
Future testing with this apparatus could potentially incorporate fluid contaminants so
sealed bearing units were selected (also effective in preventing the ingress of pin wear
debris).
37
The entire test disk/spindle assembly is driven by a DC gearmotor, mounted
horizontally to the bottom of the upper mounting plate. A 90° miter gear system is used
to transmit the power directly from the gearmotor to the spindle which provides the
relative motion between the (rotating) test disk and the static test pin (loaded from
above). The miter gear mounted on the output shaft of the gearmotor is manufactured
from case-hardened alloy steel while the mating gear (mounted to the lower shouldered
end of the spindle) is manufactured from a brass alloy. The assembly was originally
designed with a small wet sump setup below the miter gears but upon operation, it was
discovered that the higher operating speeds would cause the lubricating oil to migrate
out of the sump along the shaft of the gearmotor. It was decided that conventional shaft
seals would add to much drag to the shaft of the gearmotor so the sump was eventually
removed and a higher viscosity gear oil was used to lubricate the gear interface
(intermittent application).
Figure 15. Pin on Disk Tribometer Cross-Section View
38
The Test Pin specimens are contained in an Upper Load Arm assembly, consisting
of a number of components which are all illustrated in Figure XX. Initially, the test pins
are installed in a ’pin holder’ component (blind hole, slip fit installation) and retained
with a set screw mounted transversely. The pin holder is pinned at two locations in a
four-bar type linkage configuration which is used to ensure constant vertical alignment
of the test pin. The pin holder is pinned at the uppermost location to a control arm
assembl7y, consisting of five individual components. The control arm assembly utilizes
a RH/LH threaded rod with a threaded clevis pin at one end (mates with the pin holder)
and a female spherical bearing at the opposite end. By rotating the threaded rod, the
operator changes the effective length of the control arm assembly, which in turn changes
the angle of incidence between the pin holder and the mating disk surface. Once the
operator has dialed in the correct length of the control arm and confirmed absolute
vertical alignment of the pin holder, the length is locked down using LH/RH jam nuts.
This configuration guarantees that as the pin holder rotates about the pinned center of the
Load Arm component (to which it is also attached), the test pin will always remain
vertical and normal to the test surface.
During the initial design phase, this use of a four-bar type control arm linkage was
deemed critical to the performance of the apparatus provided the materials that were
being tested. Most often, the materials being characterized with Pin on Disk tests are
metal/metal or metal/ceramic couples and the volume loss of the pin specimen is
typically very low. In this test case, the hardness disparity between the titanium disks
and the polymer test pins is significant so the potential for wear in the softer polymer pin
is high. As the height of the test pin changes throughout the test due to wear, the angle
of incidence between the pin and the disk test surface must remain zero or the loading
profile at the test interface will consistently change (uneven bearing pressure gradient).
The Load Arm component is central to the operation of the entire test assembly.
The load arm is pinned at two locations, the outermost end is machined with a closetolerance clevis/thru hole configuration to allow installation of the pin holder and
retention with a spring pin. A transverse thru hole located approx. 1.625” from the
opposite end acts as the center of articulation for the load arm (main Pivot Pin installed
through this hole).
A larger vertical thru hole is located outboard of the pivot pin
39
location; this hole allows sufficient clearance for a threaded ‘T’ fastener which centers
through a compression spring.
Figure 16. Pin on Disk Load Arm Component
While evaluating all of the commercially available Pin-on-Disk tribometer solutions
on the market, it was noticed that in all cases, the normal force test load limit was
relatively low (max load ranging 10N to 60N). Given the range of polymer products that
will potentially be evaluated using this rig in the future, it was desired that this rig be
capable of applying FN test forces ranging 1-50+ lbf (220+ N). This capability would
allow for testing of a standard .125” (flat end) pin at effective bearing pressures up to
and exceeding 4000 psi. It was this requirement that ultimately led to the selection of a
compression spring-driven load design.
Other load application methods that were
considered during the design phase were a calibrated (slung) deadweight configuration
and also the use of hydraulic/pneumatic actuators, however the spring design offered the
best combination of adaptability, control and reliability at a significantly reduced cost.
This test assembly uses a floating ‘seat’ component made of 7075-T6511 aluminum
upon which the compression spring rests. A T-bolt runs through the center of the
compression spring (cut and ground ends). As the operator tightens the upper flange nut,
the T bolt compresses the load spring which in turn imparts an upward force on the end
of the load arm component. The upward force on the load arm creates the FN test load
on the test pin which is installed at the opposite end of the load arm. A close tolerance
40
RC fit between the Pivot Pin and the Load arm ensures that the test pin remains static
relative to the rotating test disk. Dry film lubricants are used to mitigate wear between
the Pivot Pin (manufactured from 17-4PH stainless steel) and the Load Arm while a
laminate wear pad (thermosetting resin/PTFE/Polyester fabric) is bonded to the upper
Load Plate to prevent premature wear at the contact interface when the Flange Nut is
tightened.
The inverse mechanical advantage of the load arm (4.5:1 distance from fulcrum to
test pin and applied load respectively) is used to ensure a consistent loading profile
throughout the wear regime. As the pin wears, the vertical drop has little effect on the
magnitude of the applied load (full pin sample wear results in normal force load
reduction of less than 1.1%). This extended arm geometry also provides increased
precision in the measurement of the Normal force (FN) and Drag force (FD). The Load
Arm is outfitted with four strain gauges, configured in two individual half-Wheatstone
bridge circuits to monitor the normal force applied to the test pin and also the drag force
resulting from the frictional forces between the pin and the rotating test disk. The halfbridge arrangement was used to offset thermal effects on the strain gages which are also
temperature-matched to the aluminum arm substrate (load arm manufactured from 7075T651). The ‘necked-down’ geometry of the long center section of the load arm was
designed specifically to counter the maximum design operating loads while also
allowing sufficient deflection to give accurate and precise strain readings.
41
Figure 17. Load Arm Structural Analysis at Max Anticipated Operating Load
Conditions (Spring FS=180 lbf, FD=24 lbf)
The spindle speed input which is used to calculate the relative test speed and the
total distance travelled is acquired as an analog input from an inductive-type proximity
sensor installed on the bottom of the upper mounting plate. A (stainless steel) button
head fastener is installed in a threaded hole on the bottom of the spindle; each revolution
of the test spindle passes this fastener within approximately .020” of the proximity
sensor (shown in Figure 18 below). Each time this ferrous target passes by the face of
the sensor, a +10VDC signal is passed to the data acquisition (DAQ) module; this signal
is interpreted by the HMI program and converted to test speed based on reference test
diameter:
d=
π·d·rpm/12 = VT
42
Test diameter entered by operator in
inches (.75” for this test setup)
rpm = Measured ‘Rising Edge’ events per
minute
VT =
Linear relative test sample velocity
(surface feet per minute)
Figure 18. Test Rig Proximity (Speed) Sensor
During the initial design phase, only two prime mover solutions were identified as
really being viable for this test apparatus given the desired dimensional envelope
(compact benchtop) and cost concerns.
Hydraulic and pneumatic actuators were
eliminated due to the supporting equipment required to power them (pumps with
accumulators likely required to smooth out system dynamics), leaving AC and DC
motors. The performance characteristics afforded by an alternating current (AC) motor
solution (induction-type, cap-start or servo) are well-suited to this type of application,
however the cost associated with the necessary speed controllers (VFD, servo controller,
etc.) and potential geared speed reducer proved to be prohibitive for this project. A
direct-current (DC) motor solution therefore proved to be the most cost-effective
solution for this test apparatus given the project constraints.
A 52mm 24VDC motor with an integral (planetary) gearbox was selected as the
prime mover solution for this test rig based on the performance characteristics of the
motor relative to the specific operating conditions associated with this test rig.
Consideration was provided to the duty cycles associated with the more aggressive pin
on disk tests that would be performed; the efficiency and torque/power curves for this
motor suggest that continuous duty at the upper limits of the test load range is possible.
The 17:1 reduction ratio of the planetary gearbox motor results in a 285 rpm speed rating
for this motor (IG52GM TY04), however the miter gears currently installed lend a 1.5:1
ratio which brings the actual test spindle max speed back up to around 430 rpm (actual
43
overspeeds up to 530 rpm have been achieved). Should the operator desire to run a
higher torque, lower speed test, the gear ratio is easily changed to provide additional test
capability.
FIGURE XX
A 5 amp, 0-30V DC benchtop power supply is used to energize the DC gearmotor
and drive the test disk in this apparatus. An adjustable-type power supply was selected
that offers both coarse and fine voltage adjustment controls, making it possible for the
operator to easily dial in the desired test speed within a (repeatable) resolution of
approximately ± 2 spindle rpm (±.4 sfpm for the .75” test diameter case).
All of the supporting structural components in this assembly were machined from
polymeric materials that offered sufficient strength and rigidity while also providing a
cost savings over metallic materials (such as aluminum). Ultra high molecular weight
(UHMW) polyethylene and High-density polyethylene were used as these materials also
possess reasonable chemical resistance properties (future wear testing could potentially
incorporate aggressive contaminants).
4.2 Data Acquisition System (DAQ)
Having the ability to monitor and record dynamic friction data and operating
parameters was also an important requirement for this test apparatus. Correlating realtime friction data with post-test wear results is very helpful when attempting to explain
abnormal wear phenomena. This Pin-on-Disk apparatus was designed with the ability to
monitor and dynamically record the following inputs:
44
-
Normal Force (strain)
-
Drag Force (strain)
-
Spindle RPM (proximity sensor)
As previously noted, the Load Arm component in this assembly is instrumented with
multiple strain gauges, oriented such that an accurate reading can be obtained for both
the Normal force (FN – vertical component) and Drag force (FD - horizontal component)
by measuring the deflection of the arm in the respective directions. Prior to testing, load
calibration curves were initially developed for these strain gauges using a slung
deadweight method (incremental weights).
These strain readings are recorded using a Human-Machine Interface (HMI) data
acquisition program that was developed by M. Lessard specifically for this test apparatus
using DASYLab V11 visual programming language. Figure XX below shows the actual
virtual instrument (VI) front panel seen by the operator during testing. This VI presents
the operator with real-time normal force, drag force and spindle speed data. Additional
calculations are performed dynamically in the program back plane based on inputs from
the user (reference test diameter, test pin size) and are also presented on the VI front
panel for reference: relative test speed, calculated bearing pressure, total linear distance
traveled and coefficient of friction data. All of these inputs and calculations are passed
through an averaging function in the program which writes the data to a backup file once
every ten seconds (.1 Hz) throughout the duration of the test.
then reads rising edges in this analog signal to calculate spindle RPM (
45
FIGURE XX
The HMI program for this test apparatus was designed such that each test could
be performed unattended and without interaction from the technician. Once the test
operator initiates the test sequence in the HMI program, they are prompted to enter a
Test Distance variable which is used to control an analog output trigger to a fused 5VDC
SPDT relay that controls the power supply for the gearmotor. Once the defined test
distance has been achieved, the relay is triggered (open), disconnecting the power supply
from the gearmotor and stopping the test. The double-throw relay is also wired in such a
manner that it functions as a ‘failsafe type feature, ensuring that the test will not
accumulate any cycles without recording data. In the event of any power surge event
that disables the computer, the relay releases to the ‘open’ position, effectively shutting
down the DC gearmotor.
46
5. EXPERIMENTAL RESULTS
The following sections present the findings from the metallurgical analysis
performed on each of the Ti64 thermal oxidation treatments that were conducted as well
as the results from the dynamic wear tests that were performed with each. Four different
thermal oxidation treatments were evaluated in this project (referenced as TO Case #1,
TO Case #2, etc.).
5.1 Test Disks (Thermally-Oxidized Ti64)
All of the test disk samples used in this project were manufactured using millannealed, hot-rolled 6Al-4V titanium barstock conforming to the requirements specified
in material specification AMS4928 (all material from same heat lot). Samples were
initially single-point turned from barstock in the axial direction, such that the test surface
of each sample lies in the short transverse (ST) plane with respect to the rolling direction
of the material. Both the test sample orientation and the manufacturing dimensional
requirements for each disk sample are shown in Figure XX below.
Samples machined from barstock
with orientation shown
Longitudinal Rolling
Direction (L)
Test Interface Surface
AMS4928 Barstock
All dimensions in inches
Figure XX. Test Disk Dimensional Constraints and Method of Manufacture (Barstock
Orientation)
47
After the finish turning operation, the resultant surface roughness of the upper test
interface surface on each disk was found to average around Ra 16 microinches for all
samples (measured using a Hommel Tester T1000 digital surface profilometer). The test
surface of each disk was then progressively polished using 1200# grit and 1800# grit
silicon carbide grinding paper until an average surface roughness of around Ra 2
microinches was achieved for all samples.
Figure XX shows the surface profile
inspection plot for test sample T10 prior to thermal oxidation treatment; the average
roughness (Ra) and the mean roughness (Rz) values reported for this sample were found
to be fairly representative of all samples prior to thermal treatment.
FIGURE XX. Sample # T41 – As-Polished Condition (Un-modified Titanium, No
Thermal Oxidation Treatment)
After polishing, all samples were then serialized (vibroscribed on bottom surface
T01, T02, etc.) and then sonic-cleaned in acetone. Each of the test disks were then
subjected to their respective thermal oxidation treatments. The thermal and atmospheric
controls specified in Section 3.1 for each thermal ‘case’ were strictly adhered to, using a
kiln with thermal stability of ±5°F from the nominal treatment temperature. Figure XX
depicts the condition of the test samples directly after accomplishment of each thermal
oxidation treatment.
48
Heat #1110B10
Heat #1111A10
Heat #1115A10
Heat #1113N10
Figure XX. Condition of 6Al-4V Titanium Test Disks After Each Thermal Oxidation
Treatment
After performing the various thermal oxidation treatments, each disk had a
markedly different surface appearance. The titanium disks that were subjected to the
first thermal oxidation treatment (cycle #1) maintained a dark surface layer with an
irregular ‘spotted’ appearance (lightly colored groups). These samples also showed
localized dissociation of the oxide layer at different locations (small regions of the oxide
layer appeared to have separated from the substrate during processing). The samples
subjected to thermal cycle #2 were left with a deeper rust-colored surface appearance.
These samples also showed signs of localized crazing and dissociation of the oxide layer
at different locations, however none of these samples showed signs of ‘spotting’ as the
previous samples had.
The appearance of the cycle #3 samples was significantly
different than the other three. These samples had a very light grey surface color with a
very thin and brittle dissociated oxide layer (a considerable amount of this layer flaked
away during handling, primarily from the OD surface of the disk). After treatment,
samples from cycle #4 were noticeably smoother than the other three treatment cycles
and the surface layer again had a light brown (rust-colored) appearance. The surface
49
layer of these specimens had a more normal appearance (no signs of localized
dissociation and separation of the oxide layer).
Regardless of cooling cycle, in almost every thermal oxidation treatment case
there exists small localized regions of dissociated case layer that will most likely become
liberated when a mating surface passes across it under load.
To more accurately
evaluate the tribological properties of the modified titanium case layer, all of the
thermally-oxidized samples were subjected to an additional manual polishing operation
prior to testing. This step was performed to ensure that loose surface oxides would not
impact the primary wear mechanism (potential third-body wear particles). The upper
test surface of each disk was polished using the same method as used on the pre-treated
samples, 1200# and 1800# silicon-carbide grinding papers.
In each case, material
removal was limited to .0004”-.0006” in depth to ensure that only the loose surface
oxides were removed, having the least impact on the oxygen-diffused case zone. The
resultant surface finish for all thermally-oxidized test samples after this polishing step
ranged between Ra 4and Ra 16 microinches.
Figure XX. Condition of 6Al-4V Titanium Test Disks After Thermal Oxidation
Treatment + Polishing Operation (Condition As-Tested)
50
By strictly controlling the depth of material removed from each sample, the loose
surface oxides were removed from each sample, however the dark outer TiO2 case was
not fully removed from every sample. The darker regions left on the upper test surface
of each sample (Figure XX) are indicative of relatively deep penetration by the TiO2
oxide layer. All TO-treated samples were again sonic-cleaned in acetone after this
polishing step.
5.2 Alpha Case Analysis
One sample from each of the different Thermal Oxidation treatment heat lots was
sectioned so that metallurgical (cross-section) mounts could be created and alpha case
region analyzed.
The microhardness through the case of each cross-sectioned TO
sample mount was then measured using a Knoop hardness tester (110g load), with
indications taken from .001” to .040” in depth. The hardness-depth profile for each of
the different samples is presented in Figure XX.
700
CYCLE 1
CYCLE 2
CYCLE 3
CYCLE 4
650
56.3
600
51.3
46.3
500
41.3
450
400
36.3
350
31.3
300
26.3
250
200
0.0000
21.3
0.0050
0.0100
0.0150
0.0200
0.0250
0.0300
0.0350
0.0400
Depth of Indication
Figure XX. Hardness-Depth Profile for Each Thermal Cycle (Knoop-100g Load)
51
Rockwell C (Equivalent hardness)
Knoop Hardness (HK, 100g)
550
Cross-sectional (optical) micrographs were also created for each TO sample
using a metallographic microscope so that the microstructure of each could also be
analyzed. Each of these micrographs is presented in Figure XX below. Note: due to the
increased case depth for cycle #3, this sample is presented at 200x magnification while
the remaining three samples are shown at 500x magnification.
Cycle #1 Sample
500X Mag.
Cycle #2 Sample
500X Mag.
Cycle #3 Sample
200X Mag.
Cycle #4 Sample
500X Mag.
Figure XX. Cross-Sectional Optical Micrographs Showing Microstructure Through
Alpha Case Layer in All Four TO Cycles
The total depth of affected case layer for each test sample was as follows:
Cycle #1
Cycle #2
Cycle #3
Cycle #4
0.0040 in.
0.0060 in.
0.0095 in.
0.0012 in.
52
Upon analysis, the cycle #1 sample was found to exhibit an oxygen rich case
layer with a moderate concentration of interstitial oxygen and α phase Ti (Hex ClosePacked) with decreasing gradient through a depth of approximately .0040”. A normal
transition from the TiO oxygen-diffused layer into the α+β bulk is exhibited with a fineto-coarse lamellar grain structure dominating the bulk. Throughout the TiO layer, β
(Body-Centered Cubic) Ti appears to have accumulated at prior grain boundaries in an
elongated pattern (normal to the interface surface).
The cycle #2 sample exhibited a slightly higher density of interstitial oxygen and
α phase titanium than did the cycle #1 sample. The diffusion gradient from TiO to α+β
bulk titanium was slightly more pronounced than found in the cycle #1 sample and the
resultant grain size was also slightly more coarse (again, lamellar structure). β structure
concentration found in the TiO region was slightly less, however the same elongated
grain pattern was found to be similar to that found in the cycle #1 sample. Depth of case
layer was approximately .0015” deeper than thermal cycle #1.
The affected case depth of the cycle #3 sample was significantly greater than all
of the other cases. This sample had a very dense, highly crystalline and oxygen-dense
stretching down to about .0040” in depth at which point the gradient from primary
interstitial oxygen (α) to α+β is found to occur.
It is interesting to note that this
treatment resulted in an acicular α+β structure throughout this transition region from
.0040-.0095” whereas all other samples tended to show an equiaxed lamellar structure.
The resultant grain size of the α+β region was slightly larger (more coarse) than that
found in the cycle #2 sample.
The case depth of cycle #4 was significantly lower than all other TO-treated
samples. The oxygen-rich TiO layer of this sample was only found to be approximately
.0010” in depth however the relative grain size throughout the case layer as well as the
Ti bulk was significantly finer than developed in the other three TO samples. The case
layer showed a fairly dense concentration of interstitial oxygen and α structure with a βrich region (measuring approximately .0010” in depth) found directly below. The bulk
53
titanium exhibited a fine α+β structure with β grains typically being significantly finer
than the adjacent α microstructure.
5.3 Polymer Test Pins
As required to fit the test fixture, each of the polymer test materials was
machined into a cylindrical ‘pin’ geometry having the following dimensional
constraints:
Figure XX. Polymer Pin Geometry
Being common thermoplastics that are readily available in various forms, the
UHMW polyethylene and acetal copolymer samples were machined from rod stock that
was purchased commercially (Hostalen GUR and Celcon POM respectively).
The
epoxy thermosetting two-part compound however, was cast into larger cylindrical
specimens (test tube), thermally cured, and then machined down to size (same sample
size shown in Figure XX). All specimens were manufactured using a single-point
turning procedure with no subsequent manufacturing processes.
Acetal Copolymer
Epoxy Compound
UHMW Polyethylene
Figure XX. Polymer Test Pin Samples
54
5.4 Wear Testing Results
All dynamic wear testing conducted in support of this project was performed
using a pin on-disk tribometer (described in Section 4 of this paper), adhering to the test
criteria specified in ASTM G99. Each of the four thermal oxidation treatments were
tested as well as unmodified titanium samples mated against three different polymer
materials (UHMW polyethylene, acetal copolymer and epoxy compound). The various
material couples that were tested are listed in Table X below.
Table X. Tribosystem Test Matrix
Iteration
1
2
3
4
5
6
8
9
10
11
12
13
15
16
17
19
20
21
Titanium Sample
Unmdified 6Al-4V Titanium (MA/HR)
Unmdified 6Al-4V Titanium (MA/HR)
Unmdified 6Al-4V Titanium (MA/HR)
TO Treatment 1 (Cycle #1)
TO Treatment 1 (Cycle #1)
TO Treatment 1 (Cycle #1)
TO Treatment 2 (Cycle #2)
TO Treatment 2 (Cycle #2) - Re-polished
TO Treatment 2 (Cycle #2)
TO Treatment 2 (Cycle #2) - Re-polished
TO Treatment 2 (Cycle #2)
TO Treatment 2 (Cycle #2) - Re-polished
TO Treatment 3 (Cycle #3)
TO Treatment 3 (Cycle #3)
TO Treatment 3 (Cycle #3)
TO Treatment 4 (Cycle #4)
TO Treatment 4 (Cycle #4)
TO Treatment 4 (Cycle #4)
Polymer Sample
UHMW PE
Acetal Copolymer
Epoxy
UHMW PE
Acetal Copolymer
Epoxy
UHMW PE
UHMW PE
Acetal Copolymer
Acetal Copolymer
Epoxy
Epoxy
UHMW PE
Acetal Copolymer
Epoxy
UHMW PE
Acetal Copolymer
Epoxy
Test Disk
T42
T45
T44
T01
T02
T03
T11
T17
T12
T13
T14
T15
T21
T22
T24
T31
T32
T33
Test Pin
U2
A2
E1
U3
A3
E2
U4
U7
A5
A8
E4
E7
U5
A6
E5
U6
A7
E6
Prior to testing, all titanium test disks and polymer test pin samples were
individually weighed using a Mettler Toledo JB1603-C/FACT digital scale to establish
the starting mass for each test sample. The initial (pre-test) mass of every sample was
recorded with a precision of .0001 gram weight. Based on the specific titanium wear
rates reported in other research papers, it was initially decided to use the measured
change in mass as the primary method to quantify wear rates for each test sample. Using
published wear rates, it was determined that the shallow (predicted) wear scar depth
55
could potentially lend significant variability in lost volume calculations if standard
profilometers were used to gauge depth.
To allow for comparison against published results, it was desired that the
operating speeds, pressures (etc) used in this series of testing be very close to those most
commonly used in historical testing with the UHMW/TO titanium tribosystem. Initial
calibration tests were therefore performed using the operating parameters specified in the
work published by H. Dong [X] (5 MPa/.25 m/s = 725 psi/50 sfpm), however these
parameters were found to be too aggressive for these polymer materials in the absence of
water lubrication. Additional calibration tests were performed using a normal load equal
to 50% of the pressure rating for the acetal and UHMW materials with a reduced speed.
Preliminary wear rates were found to be reasonable and these operating parameters were
then used for all test iterations (all material combinations tested using the same
load/speed):
Table XX. Pin on Disk Testing Parameters
Test Load (FN applied to pin)
Calculated Effective Bearing Pressure
Reference Wear Track Diameter
Spindle Speed
Relative Test Speed
Minimum Test Distance1
15
543
0.75
153
30
lbf
psi
in.
rpm
sfpm
40000 ft
1
Individual test iterations were allowed to continue until either the
minimum test distance was achieved or the polymer test pin showed
signs of excessive wear. Some samples were tested beyond the
40,000 ft minimum distance.
All pin-on-disk tests were performed sequentially (according to the order shown
in Table X) with the exception of iterations 9, 11 and 13. The initial tests performed
with the Cycle #2 samples yielded wear rates that were significantly higher than
expected (see Tables X, X and X). Reviewing the pre-test surface condition on the Cycle
#2 samples revealed a significant concentration of residual surface oxides and it is
conceivable that this condition could have a negative impact on the titanium wear rates
developed during testing. If the strength of adhesion between this residual oxide layer
56
and the TiO substrate is relatively low, it is possible that the TiO2 oxides could become
liberated during testing, embedding in the plastic counterface and causing accelerated Ti
wear rates.
To assess the impact that this surface condition might have on the wear
mechanisms in this tribosystem, three additional samples were added to the testing
regimen (iterations 9, 11 and 13). These added Cycle #2 samples were polished a
second time prior to testing using the same process (1200#/1800# progressive polish
using SiC papers), however each of these samples was polished until all evidence of the
penetrated oxide layer was removed. This resulted in an additional .0010’-.0015” of
material removal (still within the diffused TiO zone of the sample). These three samples
were then tested using the same operating conditions and mating materials.
As noted previously, the actual mass of every test sample was measured prior to
testing and recorded with a precision of .0001 grams. After completion of the pin on
disk test, the titanium disk samples were sonic cleaned in acetone, thoroughly dried
using pressurized air and also allowed to dry in air for 10 minutes. Each sample was
then weighed and titanium wear rates were calculated based on the difference in weight
(volume calculated using a density of .160 lb/in3 for 6Al-4V). Pin samples were wiped
clean (no solvents or liquids) and also weighed. Wear rates were again calculated based
on change in mass (material density used for UHMW, acetal copolymer and epoxy are
.036, .050 and .053 lb/in3 respectively).
57
Figure X. Pin-on–Disk Tribometer
Titanium-UHMW Testing
Iteration
1
4
8
9
15
19
Titanium Sample
Unmodified 6Al-4V Titanium (MA/HR)
TO Treatment 1 (Cycle #1)
TO Treatment 2 (Cycle #2)
TO Treatment 2 (Cycle #2) - Re-polished
TO Treatment 3 (Cycle #3)
TO Treatment 4 (Cycle #4)
Polymer Sample
UHMW PE
UHMW PE
UHMW PE
UHMW PE
UHMW PE
UHMW PE
Test Disk
T42
T01
T11
T17
T21
T31
Test Pin
U2
U3
U4
U7
U5
U6
The final results of all wear tests performed using UHMW polyethylene as a
mating countersurface are shown in Figure XX below; results are presented in terms of
specific titanium wear (total wear divided by total distance accumulated and test load).
The IF values listed for each test convey the calculated improvement factor, or the ratio
of wear between that specific test and the wear rate for the unmodified titanium baseline.
For example, an IF factor of 5.0 indicates a specific wear rate that was 5x lower than that
of the baseline sample and an IF factor of 0.5 would indicate a wear rate that was 2x
higher than the baseline.
58
BASELINE WEAR RATE (Unmodified Titanium)
Test Disk Sample T42 / UHMW Sample U2
5.27·10-11 in3/lb·ft
6
IF = 3.3
5
THERMAL CYCLE #1
Test Disk Sample T01 / UHMW Sample U3
1.61·10-11 in3/lb·ft
4
THERMAL CYCLE #2
Test Disk Sample T11 / UHMW Sample U4
9.29·10-11 in3/lb·ft
3
THERMAL CYCLE #2 (2) - Repolished Sample*
Test Disk Sample T17 / UHMW Sample U7
6.86·10-11 in3/lb·ft
THERMAL CYCLE #3
Test Disk Sample T21 / UHMW Sample U5
1.61·10-12 in3/lb·ft
2
THERMAL CYCLE #4
Test Disk Sample T31 / UHMW Sample U6
3.18·10-12 in3/lb·ft
1
0
1
2
3
IF = 0.6
IF = 0.8
IF = 32.7
IF = 16.6
4
5
6
7
8
9
10
Titanium Disk Specific Wear Rate (x 10 -11 in3/lb·ft)
Figure XX. Titanium Specific Wear Rate Comparison - Titanium Test Disks Tested
Against UHMW Polyethylene (530 psi / 30 sfpm)
It is interesting to note that the Cycle #2 sample actually developed a higher wear
rate than the baseline titanium.
Additional polishing of the Cycle #2 test surface
improved the wear rate slightly (however, still higher than the unmodified Ti64 sample).
After completion of the wear test, the unmodified titanium sample was found to exhibit
the typical signs of wear distress (grooved scarring) across the test interface. All other
(thermally-oxidized) samples showed evidence of contact and generally mild wear; none
showed evidence of the deep wear grooving displayed on the unmodified Ti64 sample
test surface (including the Cycle #2 sample).
59
Sample T42 (Baseline)
Sample T42
Sample T21
Sample T01
Sample T11
Sample T31
Figure XX.
Real-time dynamic load and friction data was also recorded for all test iterations
using the integral data acquisition capabilities of the test apparatus. Figure XX shows
the friction values recorded for each of the UHMW-mated samples as a function of
distance traveled (and with test velocity set to a constant speed, the plot also correlates
linearly to elapsed test time).
60
0.5
1
COF Vs. DISTANCE TRAVELED
0.9
0.45
0.4
Ti64 TEST DISKS Vs.
UHMW POLYETHYLENE PIN
0.35
0.3
0.8
0.25
0.2
0.15
COEFFICIENT OF FRICTION (μ)
0.7
0.1
0.05
0.6
0
0
0.5
2000
4000
6000
8000
10000
12000
14000
Data Acqusition System Disconnect
(Test Sample T31)
0.4
0.3
0.2
Data Acqusition System Disconnect
(Test Sample T21)
T42 - UN-MODIFIED TITANIUM (MAHR-AMS4928)
T01 - THERMAL CYCLE (1)
T11 - THERMAL CYCLE (2)
T21 - THERMAL CYCLE (3)
T31 - THERMAL CYCLE (4)
0.1
0
0
20000
40000
60000
80000
100000
120000
140000
LINEAR DISTANCE TRAVELED (FT)
Figure XX. Coefficient of Friction Data Vs. Distance Traveled (Constant Velocity) for
all Ti64 – UHMW Test Iterations
During testing with samples T21 and T31, it was decided to allow these samples
to accumulate a greater total distance of travel (110,000+ linear ft versus the min
distance of 40,000 for other samples) to allow verification of friction trends. The final
specific wear rates are reported as a function of both distance and load (weight loss per
unit distance and unit force) so all wear results are comparable regardless of total
distance traveled. At different times during each of these test iterations however, the
data acquisition system experienced a software disconnect and no data was recorded
(while testing continued at constant velocity and load).
This gap in COF data is
reflected in Figure X and it should be noted that this DAQ malfunction had no impact on
the reported wear rate results.
61
The specific wear rates for each of the UHMW polymer pin samples in this series
was also calculated at the end of each test based on the change in mass. These samples
were again measured before testing and also at the end of each test using a Mettler
Toledo JB1603-C/FACT digital scale. Figure XX compares the wear rate for each pin
sample against the baseline test and also provides the ‘Improvement Factor’ for each
case.
BASELINE WEAR RATE (Unmodified Titanium)
Test Disk Sample T42 / UHMW Sample U2
2.34·10-10 in3/lb·ft
6
THERMAL CYCLE #1
Test Disk Sample T01 / UHMW Sample U3
0.51·10-10 in3/lb·ft
5
IF = 4.6
4
THERMAL CYCLE #2
Test Disk Sample T11 / UHMW Sample U4
1.32·10-10 in3/lb·ft
3
THERMAL CYCLE #2 (2) - Repolished Sample*
Test Disk Sample T17 / UHMW Sample U7
0.97·10-10 in3/lb·ft
IF = 2.4
THERMAL CYCLE #3
Test Disk Sample T21 / UHMW Sample U5
0.47·10-10 in3/lb·ft
2
IF = 5.0
1
THERMAL CYCLE #4
Test Disk Sample T31 / UHMW Sample U6
1.20·10-10 in3/lb·ft
0.000
IF = 1.8
0.500
IF = 1.9
1.000
1.500
UHMW Pin Specific Wear Rate (x 10-10
2.000
2.500
in3/lb·ft)
Figure XX. UHMW Specific Wear Rate Comparison - UHMW Test Pins Tested
Against Various Titanium Test Disks (530 psi / 30 sfpm)
Titanium-Acetal Copolymer Testing
Iteration
2
5
10
11
16
20
Titanium Sample
Unmodified 6Al-4V Titanium (MA/HR)
TO Treatment 1 (Cycle #1)
TO Treatment 2 (Cycle #2)
TO Treatment 2 (Cycle #2) - Re-polished
TO Treatment 3 (Cycle #3)
TO Treatment 4 (Cycle #4)
62
Polymer Sample
Acetal Copolymer
Acetal Copolymer
Acetal Copolymer
Acetal Copolymer
Acetal Copolymer
Acetal Copolymer
Test Disk
T45
T02
T12
T13
T22
T32
Test Pin
A2
A3
A5
A8
A6
A7
All of the titanium disk samples that were tested against acetal copolymer again
showed an improvement in wear rates as compared against the baseline; with the
exception of cycle #2 (as originally processed).
After the initial sequence of test
iterations was completed, a second Cycle #2 sample was tested again with a re-polished
test surface as discussed above. For this tribosystem, the additional polishing step
(complete removal of residual surface oxides) significantly improved the wear
performance, knocking the wear rate down from 133% of the baseline to 29% of the
baseline (or an improvement factor of 3.4).
BASELINE WEAR RATE (Unmodified Titanium)
Test Disk Sample T45 / ACETAL Sample A2
4.04·10-11 in3/lb·ft
6
THERMAL CYCLE #1
Test Disk Sample T02 / ACETAL Sample A3
1.66·10-11 in3/lb·ft
5
4
THERMAL CYCLE #2
Test Disk Sample T12 / ACETAL Sample A5
5.38·10-11 in3/lb·ft
3
THERMAL CYCLE #2 (2) - Repolished Sample*
Test Disk Sample T13 / ACETAL Sample A8
1.17·10-11 in3/lb·ft
THERMAL CYCLE #3
Test Disk Sample T22 / ACETAL Sample A6
6.76·10-12 in3/lb·ft
2
THERMAL CYCLE #4
Test Disk Sample T32 / ACETAL Sample A7
NO MEASURABLE WEAR
1
0
1
2
IF = 2.4
IF = 0.7
IF = 3.4
IF = 6.0
IF = INF.
3
4
5
6
Titanium Disk Specific Wear Rate (x 10 -11 in3/lb·ft)
Figure XX. Titanium Specific Wear Rate Comparison - Titanium Test Disks Tested
Against UHMW Polyethylene (530 psi / 30 sfpm)
Probably the most notable test iteration from the Ti-acetal tests, however was the
TO Cycle #4 sample. After accumulating nearly 44,000 ft of distance traveled using the
same load and speed parameters as was used in all other tests, this sample exhibited no
measureable wear (same pre- and post-test mass out to .0001 grams). After testing, this
63
sample exhibited a very faint burnished appearance in the contact region with no signs of
wear distress.
Sample T45 (Baseline)
Sample T45
Sample T22
Sample T02
Sample T12
Sample T32
Figure XX.
After testing, the unmodified 6Al-4V titanium sample again showed the same
type of distress mode at the contact interface (rough wear scarring). Similar to the
results witnessed with the UHMW test iterations, none of the thermally oxidized samples
tested in this series exhibited this type of distress after testing (no detectable wear scars).
To varying degrees, however, each of these samples showed slight signs of burnishing
(polished appearance) in the test contact zone.
Plots showing the coefficient of friction data for each of these test samples can be
found in Figure XX. During testing with the Cycle #3 sample (T22), the technician
inadvertently disabled power to the test rig drive motor, bringing the test speed down to
zero. Power was then immediately restored but the test speed was set to 40 surface feet
per minute (sfpm) rather than the desired 30 sfpm. After the sample had accumulated
approximately 2000 feet at this speed, it was then recognized and reduced to 30 sfpm.
Interestingly, the inadvertent increase in test speed actually resulted in a much higher,
non-linear coefficient of fiction during this time (this condition is identified in Figure
XX below).
64
0.5
1
COF Vs. DISTANCE TRAVELED
0.9
Ti64 TEST DISKS Vs.
ACETAL COPOLYMER PIN
0.8
0.45
0.4
0.35
0.3
0.25
0.2
0.15
T22 Restart speed
initially set at 40 SFPM
0.1
then reduced to 30 SFPM
COEFFICIENT OF FRICTION (μ)
0.7
0.05
0.6
0
0
1000
2000
3000
4000
5000
6000
7000
8000
0.5
0.4
0.3
0.2
0.1
T45 - UN-MODIFIED TITANIUM (MAHR-AMS4928)
T02 - THERMAL CYCLE (1)
T12 - THERMAL CYCLE (2)
T22 - THERMAL CYCLE (3)
T32 - THERMAL CYCLE (4)
Test Rig Shutdown / Restart
(Test Sample T22)
0
0
10000
20000
30000
40000
50000
60000
70000
LINEAR DISTANCE TRAVELED (FT)
Figure XX. Coefficient of Friction Data Vs. Distance Traveled (Constant Velocity) for
all Ti64 – Acetal Copolymer Test Iterations
Each of the polymer test pins tested in this series were also measured prior-to and
following completion of each test so that specific wear rates could be calculated for
these components. Figure XX provides a comparison between the acetal copolymer
wear rates as mated against the baseline titanium system and also the individual
thermally-oxidized Ti samples. The thermal cycle #1 sample was found to exhibit the
lowest wear rates of all tested samples by far (about 2.4% of the baseline wear rate).
This correlates well with the low, normal COF values recorded during testing.
65
BASELINE WEAR RATE (Unmodified Titanium)
Test Disk Sample T45 / ACETAL Sample A2
11.37·10-10 in3/lb·ft
6
5
THERMAL CYCLE #1
Test Disk Sample T02 / ACETAL Sample A3
0.28·10-10 in3/lb·ft
4
THERMAL CYCLE #2
Test Disk Sample T12 / ACETAL Sample A5
12.98·10-10 in3/lb·ft
3
THERMAL CYCLE #2 (2) - Repolished Sample*
Test Disk Sample T13 / ACETAL Sample A8
6.73·10-10 in3/lb·ft
2
THERMAL CYCLE #3
Test Disk Sample T22 / ACETAL Sample A6
10.07·10-10 in3/lb·ft
IF = 0.9
IF = 1.7
IF = 5.0
THERMAL CYCLE #4
Test Disk Sample T32 / ACETAL Sample A7
IF = 1.9
1
IF = 41.1
4.76·10-10 in3/lb·ft
0.000
2.000
4.000
6.000
8.000
10.000
12.000
14.000
Acetal Pin Specific Wear Rate (x 10-10 in3/lb·ft)
Figure XX. UHMW Specific Wear Rate Comparison – Acetal Copolymer Test Pins
Tested Against Various Titanium Test Disks (530 psi / 30 sfpm)
Titanium-Epoxy Testing
Iteration
3
6
12
13
17
21
Titanium Sample
Unmodified 6Al-4V Titanium (MA/HR)
TO Treatment 1 (Cycle #1)
TO Treatment 2 (Cycle #2)
TO Treatment 2 (Cycle #2) - Re-polished
TO Treatment 3 (Cycle #3)
TO Treatment 4 (Cycle #4)
Polymer Sample
Epoxy
Epoxy
Epoxy
Epoxy
Epoxy
Epoxy
Test Disk
T44
T03
T14
T15
T24
T33
Test Pin
E1
E2
E4
E7
E5
E6
Each of the titanium disk test disk samples that were tested against the epoxy
compound bulk material displayed signs of wear distress on the titanium test surface
during testing.
The unmodified (baseline) test disk sample developed a very
pronounced, rough wear scar on the upper test surface with significant volume loss in a
relatively short distance traveled (4300 feet). Each of the thermally-oxidized titanium
66
samples also showed signs wear distress and burnishing at the test interface following
testing, however not nearly to the extent that was witnessed with the baseline test
sample. Figure XX shows the post-test condition of three of the epoxy-mated test
samples. Note: These samples were sectioned after testing and subjected to separate
analysis not directly related to this project.
Sample T44 (Baseline)
Sample T44
Sample T03
Sample T24
Sample T14
Sample T33
Figure XX.
Each of the tests conducted using epoxy compound were limited in distance by
the wear rate of the actual polymer test pin; the wear rate for this material was found to
be significantly higher than experienced with the UHMW and acetal mating materials.
Each test iteration was allowed to continue until the majority of the pin material had
worn away, tests were stopped when the pin-holder component came within close
proximity of the disk test face (all tests stopped prior to contact). During the baseline
test iteration (unmodified titanium), the rate of epoxy wear was such that it was
necessary to stop the test after 4,300 feet of distance traveled.
Testing with the
thermally-oxidized test samples resulted in lower specific wear rates and these tests were
allowed to continue past 10,000 ft (Cycle #4 tested to 17,000 ft).
67
Figure XX shows an image of sample T24 that was taken during the actual test.
The brown-colored particles and fibers seen around the base of the sample and bordering
the contact path is wear debris from the epoxy material. Unlike the UHMW and acetal
materials, as the test progressed the epoxy samples generated a fair amount of this brown
wear debris (acetal and UHMW samples developed small amounts of ‘stringy’
white/grey debris).
Test disk
Test pin
Small, liberated
fragments of epoxy pin
Wear Debris
(epoxy)
Figure XX. Test Sample T24 (Cycle #3) Tested Against Epoxy Compound (Image
Captured Mid-Test Using Test Rig Remote Macro Cam)
Each test conducted with the epoxy mating material was also characterized by a
constant, audible ‘squealing’ noise. Small fragments of the test pin were also noticed
around the base of the test sample after completion (fractured/liberated during testing).
It is believed that the audible noise emitted during each epoxy test is indicative of a
stick-slip mechanism that is occurring at a very high frequency and on a micro scale.
This condition could be contributing to the fracture distress that was witnessed during
these tests. As these fractured pieces were not removed by the traditional wearing
mechanism resulting from the dynamic face-face contact, they were also included in the
final weight of the test pin.
68
The specific wear rate results for each of the titanium disk components tested in
this series (epoxy mating surface) is provided in Figure XX. It was during these tests
that the greatest reduction in wear of both the test disk and the test pin was realized when
comparing the thermally-oxidized samples against the baseline. In all cases, the rate of
titanium wear (disk) was at least an order of magnitude lower with the TO samples than
was developed during the baseline test. Testing with the Cycle #4 sample resulted in a
wear rate that was almost 3 orders of magnitude lower than the baseline (the change in
mass of the test disk over the course of the 16,300 ft test was only .0001 grams).
BASELINE WEAR RATE (Unmodified Titanium)
Test Disk Sample T44 / EPOXY Sample E1
5.00·10-9 in3/lb·ft
6
THERMAL CYCLE #1
Test Disk Sample T03 / EPOXY Sample E2
3.58·10-10 in3/lb·ft
5
IF = 14.0
4
THERMAL CYCLE #2
Test Disk Sample T14 / EPOXY Sample E4
3.13·10-10 in3/lb·ft
IF = 16.0
3
THERMAL CYCLE #2 (2) - Repolished Sample*
Test Disk Sample T15 / EPOXY Sample E7
3.13·10-10 in3/lb·ft
IF = 16.0
THERMAL CYCLE #3
Test Disk Sample T24 / EPOXY Sample E5
2.29·10-10 in3/lb·ft
2
THERMAL CYCLE #4
Test Disk Sample T33 / EPOXY Sample E6
5.78·10-10 in3/lb·ft
1
0
100
200
IF = 21.8
IF = 864.3
300
Titanium Disk Specific Wear Rate (x 10
400
-11
500
600
3
in /lb·ft)
Figure XX. Titanium Specific Wear Rate Comparison - Titanium Test Disks Tested
Against Epoxy Compound (530 psi / 30 sfpm)
The coefficient of friction data recorded for all epoxy-mated test samples is
shown in Figure XX. In all tests conducted with an epoxy mating surface, the friction
69
values were found to be fairly high (0.6-0.85) and also very unstable (constant
fluctuations up/down).
1
T44 - UN-MODIFIED TITANIUM (MAHR-AMS4928)
T03 - THERMAL CYCLE (1)
T15 - THERMAL CYCLE (2)
T24 - THERMAL CYCLE (3)
T33 - THERMAL CYCLE (4)
COF Vs. DISTANCE TRAVELED
Ti64 TEST DISKS Vs.
EPOXY COMPOUND PIN
0.9
0.8
COEFFICIENT OF FRICTION (μ)
0.7
0.6
1
0.5
0.9
0.8
0.4
0.7
0.6
0.3
0.5
0.4
0.3
0.2
0.2
0.1
0.1
0
0
500
1000
1500
2000
0
0
2000
4000
6000
8000
10000
12000
14000
16000
18000
LINEAR DISTANCE TRAVELED (FT)
Figure XX. Coefficient of Friction Data Vs. Distance Traveled (Constant Velocity) for
all Ti64 – Epoxy Compound Test Iterations
The temperature of the test disk component was also monitored manually during
each test using an exposed thermocouple method and the actual heat generated during
each test using epoxy was found to be anywhere from 30%-100% higher than was
witnessed with the UHMW and acetal samples. Rough magnitude temperatures ranged
from around 120F to 180F for the UHMW / Acetal tests and around 200F-350F for the
epoxy tests. Note: these temperatures do not reflect the actual temperature at the test
interface. Location of thermocouple tip varied from test to test – close standoff side of
disk, bottom of disk, etc.
70
Figure XX presents the specific wear rates that were measured for each of the
epoxy pin components after testing. Similar to the titanium disk wear rates in the
respective tests, all of the thermally-oxidized test samples showed a significant
improvement over the baseline rate in terms of polymer wear. The Cycle #4 sample
again showed the greatest reduction in wear (approximately 48 times lower than the
baseline sample).
BASELINE WEAR RATE (Unmodified Titanium)
Test Disk Sample T44 / EPOXY Sample E1
620.8·10-10 in3/lb·ft
6
5
THERMAL CYCLE #1
Test Disk Sample T03 / EPOXY Sample E2
75.5·10-10 in3/lb·ft
IF = 8.2
4
THERMAL CYCLE #2
Test Disk Sample T14 / EPOXY Sample E4
73.4·10-10 in3/lb·ft
IF = 8.5
3
THERMAL CYCLE #2 (2) - Repolished Sample*
Test Disk Sample T15 / EPOXY Sample E7
132.3·10-10 in3/lb·ft
IF = 4.7
2
THERMAL CYCLE #3
Test Disk Sample T24 / EPOXY Sample E5
152.9·10-10 in3/lb·ft
IF = 4.1
THERMAL CYCLE #4
Test Disk Sample T33 / EPOXY Sample E6
12.9·10-10 in3/lb·ft
1
0
100
200
IF = 48.3
300
400
Epoxy Pin Specific Wear Rate (x 10-10
500
600
700
in3/lb·ft)
Figure XX. UHMW Specific Wear Rate Comparison – Epoxy Compound Test Pins
Tested Against Various Titanium Test Disks (530 psi / 30 sfpm)
71
6. DISCUSSION AND CONCLUSIONS
Comparing all of the wear results presented herein, it is evident that the application
of thermal oxidation surface engineering techniques to titanium alloys can significantly
improve the wear and friction properties of these materials with respect to polymer
mating materials. Testing has shown that both titanium and polymer wear rates can be
reduced by several orders of magnitude in these tribosystems by introducing an oxygenrich, hardened case layer on the surface of the titanium component and carefully
controlling the physical characteristics of this surface. This engineered surface has
proven successful in offsetting the abnormal wear mechanism that is known to exist in
many titanium-polymer tribosystems.
The theoretical background for this project centers on the notion that the abnormal
wear phenomenon developed in dynamic titanium-polymer material systems is most
likely caused by the breakdown of certain hydrogen-containing polymer groups during
operation. This reaction then fuels a hydrogen embrittlement mechanism in the titanium
component which reduces the toughness of the bulk material, ultimately causing
aggressive wear of the titanium material by the (much softer) plastic mating surface.
Historically [7, 8], researchers have been able to prove this theory in dynamic UHMWTi64 material couples by analyzing titanium test surfaces for the presence of hydrides
after testing. The discovery of titanium hydrides formed below the interface surface
appears to be a fairly conclusive finding, supporting the theory of hydrogen-assisted
wear in the UHMW polyethylene – 6Al-4V titanium system.
It is important to note that the ultimate goal of this project was strictly to evaluate
the tribological performance of titanium materials mated with three specific engineering
polymers, and also the improvements that can be achieved by applying different types of
thermal oxidation treatments to the titanium interface surface. Detecting hydrides in
titanium bulk materials typically requires very specialized equipment such as a
transmission electron microscope (TEM) due to the extremely small size of these
elements. This resource was not available and looking for the presence of hydrides was
72
outside the scope of this project. The polymer materials being evaluated in this project,
however have been identified as causing the same type of wear distress on titanium
interface surfaces that was witnessed during UHMW testing and the theory of hydrogenassisted wear offers a reasonable explanation for this phenomenon.
The initial baseline testing that was performed using mill-annealed, hot-rolled 6Al4V (unmodified) titanium produced good results, generating wear distress of the
titanium interface surface that was very similar to that described in historical testing.
Examination after testing showed evidence of deep wear scarring with localized regions
showing a ‘pitting’-type wear (liberated particles). Figure XX shows a location in the
wear scar of titanium sample T42 (unmodified Ti) after testing against UHMW
polyethylene at 530 psi / 30 sfpm (41,842 ft total distance traveled).
Polymer (UHMW)
material transfer
Pronounced ‘grooving’
of test surface
Deep scarring of the
Titanium interface surface
Small ‘pitting’ zones –
liberated material
Figure X. T42 Wear Scar Morphology (50x magnification)
Notes:
Case #1
73
Case #2
Case #3
-
Relative oxygen density versus hardness
-
Appearance = grey, no detectable TiO2 layer (as expected)
Case #4
-
Hardness-Depth Profile does not coincide with published results; due to furnace
cool? (Not the case, those parts also FC)
-
Also, microhardness not indicative of actual surface hardness due to .0010” depth
measurement
-
No spallation, highly adherent oxide scale – primarily rutile, some anatase
-
Much finer grain structure
-
More surface uniformity
74
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