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ROSEMAR BATISTA DA SILVA
PERFORMANCE OF DIFFERENT CUTTING TOOL
MATERIALS IN FINISH TURNING OF Ti-6Al-4V
ALLOY WITH HIGH PRESSURE COOLANT SUPPLY
TECHNOLOGY
UNIVERSIDADE FEDERAL DE UBERLÂNDIA
FACULDADE DE ENGENHARIA MECÂNICA
2006
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Dados Internacionais de Catalogação na Publicação (CIP)
S586p
Silva, Rosemar Batista, 1974-
Performance of different cutting tool materials in finish turning of
Ti-6Al-4V alloy with high pressure coolant supply / Rosemar Batista
Silva. - 2006.
299 f. : il.
Orientadores: Alisson Rocha Machado, Emmanuel Okechukwu Ezugwu.
Tese (doutorado) - Universidade Federal de Uberlândia, Programa de
Pós-Graduação em Engenharia Mecânica; London South Bank University.
Inclui bibliografia.
1. Usinagem - Teses. 2. Liga de titânio - Teses. 3. Fluidos de corte -
Teses. I. Machado, Alisson Rocha. II. Ezugwu, Emmanuel Okechukwu.
III. Universidade Federal de Uberlândia. Programa de Pós-Graduação em
Engenharia Mecânica. IV. London South Bank University. V. Título.
CDU: 621.9
Elaborada pelo Sistema de Bibliotecas da UFU / Setor de Catalogação e Classificação
ads:
ROSEMAR BATISTA DA SILVA
PERFORMANCE OF DIFFERENT CUTTING TOOL
MATERIALS IN FINISH TURNING OF Ti-6Al-4V ALLOY
WITH HIGH PRESSURE COOLANT SUPPLY
TECHNOLOGY
Tese apresentada ao Programa de Pós-
Graduação em Engenharia Mecânica da
Universidade Federal de Uberlândia -
modalidade Sanduíche no Exterior realizado com
a London South Bank University – Londres,
Reino Unido, como parte dos requisitos para a
obtenção do título de DOUTOR EM
ENGENHARIA MECÂNICA.
Área de Concentração: Materiais e Processos de
Fabricação
Orientadores:
- Prof. Dr. Alisson Rocha Machado (Brasil)
- Prof. Dr. Emmanuel Okechukwu Ezugwu
(Inglaterra)
UBERLÂNDIA – MG
2006
iii
To my parents, Mrs. Rosalina Batista and Mr. Antônio Caetano,
and sisters
Ms. Rejaine Inês, Ms. Jeane Inês, Ms. Josiane Inês
for their encouragements and love.
To the memory of Mr. Jovino Batista da Fonseca,
my unforgettable grandfather.
iv
DECLARATION
The research presented in this thesis is the original work of the author except where
otherwise specified by references, or where acknowledgements are made. The project was
carried out at the Machining Research Centre, Faculty of Engineering Science and the Built
Environment (England), and Laboratory of Teaching and Research in Machining, Faculty of
Mechanical Engineering of Federal University of Uberlândia (Brazil) under the supervision of
Professor E. O. Ezugwu and Professor A.R. Machado, respectively. This work is being
submitted for the degree of Doctor of Philosophy – Ph.D. jointly to London South Bank
University (UK) and Universidade Federal de Uberlândia – Brazil.
v
ACKNOWLEDGEMENTS
The author, R.B. Da Silva, wishes to thank:
• Professor E.O. Ezugwu (my British Supervisor) for his magnificent supervision,
encouragement, enormous patience, support, specialised advice, professionalism and
constructive suggestions throughout the developing of this research work and for carefully
reviewing the manuscripts.
• Professor A. R. Machado (my Brazilian Supervisor), a special person to whom I am in
debt for the rest of my life because he believed in myself and gave me the opportunity to
know and work with Prof. Ezugwu. His dedication, excellent supervision, professionalism,
guidance, encouragement and contributions will be always remembered.
• Dr. J. Bonney, a friend and colleague who I also consider my Supervisor, for his patience,
assistance, invaluable help on machining trials and valuable suggestions throughout the
development of this research work.
• Conselho Nacional de Desenvolvimento Científico e Tecnológico – CNPq and Instituto
Fábrica do Milênio, both in Brazil, for their financial support throughout the course of this
research work.
• Rolls-Royce plc for funding this study providing workpiece materials and the cutting tools
employed for the research project. Sincere thanks should go to Mr. I. Baker and Mr. J.
Watkins (retired) both from Rolls-Royce, whose superb project management and leadership
skills as well as persistent encouragements led to successful conclusion of the JSF GAF
project.
• Mr. A. R. Shabbazz Nelson, a dear friend whose knowledge is greatly admired, for his
help in English Language study and invaluable advices.
• Technical staff of the Faculty of Engineering, Science and the Built Environment,
especially to Mr. J. Heyndyk and Mr. B. Hopday-Pepper who contributed to this work.
Sincere thanks to Mr. W. Winter who prepared the chemical etchants for Ti-6Al-4V alloy
workpiece samples.
• Postgraduate Programme of Faculdade de Engenharia Mecânica da Universidade Federal
de Uberlândia for providing the necessary research facilities and allowing me to go this thesis.
• My Brazilian colleagues from Faculdade de Engenharia Mecânica (FEMEC-UFU) in
particular Dr. M.B da Silva, Dr. A.M. Reis, Dr. E.S. Costa, Mr. F.J. da Silva, Mr. F. Neto, Mr.
vi
P.R. Mota, Mr. U.B. Souto, Ms. D. O. Almeida, Mr. I. L. Siqueira, Mr. N. E. Luiz and Mr. R.
Viana for their help, consideration and friendship.
• Finally my parents, Mrs. Rosalina Batista and Mr. Antônio Caetano, and sisters Ms.
Rejaine Inês, Ms. Jeane Inês for their prayers, encouragements, support and trust. In particular
I am very grateful to my youngest sister Ms. Josiane Inês who left her activities in Brazil to
help and support me in London.
vii
Da Silva, R.B. Performance of Different Cutting Tool Materials in Finish Turning of Ti-6Al-4V Alloy
with High Pressure Coolant Supply Technology, 2006, 299 f. Ph.D. Thesis, Universidade Federal de
Uberlândia, Uberlândia.
ABSTRACT
This study investigated the machinability of Ti-6Al-4V alloy with newly developed cutting tools such
as uncoated (T1 and T3) and coated (T2 and T4) cemented carbides, Polycrystalline Diamond (PCD) –
T5 and T6 inserts, Cubic Boron Nitride (CBN) – T7,T8,T9 inserts, SiC Whiskers Reinforced Ceramic
(T10) insert, and Al
2
O
3
base (T11) and Si
3
N
4
base nano-grain size ceramic (T12) inserts using various
cooling environments such as high pressure coolant supplies at pressures of 7 MPa, 11 MPa and
20.3 MPa, argon enriched environment and conventional coolant flow at high speed machining
conditions typical of finish turning operation. Tool life and failure modes, wear mechanisms,
component forces generated, surface integrity, surface finish and chip form data were used to assess
the performance of the different cutting tools and cooling environments investigated. PCD and carbide
inserts gave the best performance, in terms of tool life, when machining Ti-6Al-4V alloy. In general
coarser (T1 and T4) grain size carbides and PCD (T5) inserts gave the best overall performance in
terms of lower wear rate hence longer tool life compared to finer grain (T2,T3 and T6) grades.
Encouraging tool life can be achieved when machining with high pressure coolant supply relative to
conventional coolant flow and in the presence of argon. Tool lives generally increased with increasing
coolant pressure due to the ability of the high coolant pressure to reduce the tool-chip contact
length/area and to lift the chip, thereby providing adequate lubrication at the tool-chip interface with
consequent reduction in friction. Machining with T1, T4 and T10 inserts in presence of argon was only
able to prevent chip ignition with no improvement in tool life, due probably to the suppression of the
cooling and/or lubrication characteristics of argon gas when machining at cutting conditions
investigated. Up to 8 fold improvement in tool life were achieved when machining with PCD inserts
relative to carbide inserts under conventional coolant flow. All the grades of CBN inserts gave poor
performance during machining due to accelerated nose wear and, in some cases, severe chipping of the
cutting edge associated with a relatively high diffusion wear rate that tends to weaken the bond
strength of the tool substrate. An increase in the CBN content tends to accelerate notch wear rate,
consequently diminishing tool life under the cutting conditions investigated. Micron and nano-grain
size ceramics did not demonstrate satisfactory performance in terms of tool wear rate and tool life, due
to severe abrasive wear and chipping of the cutting edge, hence the poor machined surfaces generated.
Nose wear was the dominating tool failure mode when machining with carbide, PCD and CBN (T7)
inserts due to a reduction in tool-chip and tool-workpiece contact lengths and the consequent increase
in both normal and shear stresses and temperature at the tool tip, while severe notching and chipping
occurred when machining with CBN (T8 and T9) and micron grain size ceramics. Severe notching
also occurred when machining with nano-grain ceramic inserts, often leading to catastrophic tool
failure at speeds in excess of 110 m min
-1
. Machining with PCD tools gave lower cutting forces than
carbides inserts. Surface roughness values generated with carbides, PCD and CBN inserts were
generally within the 1.6 µm rejection criterion for finish machining and above 2 µm when machining
with all grades of ceramics employed. Micrographs of the machined surfaces show that micro-pits are
the main damage to the machined surfaces. Microhardness of the machined surfaces when machining
with carbides varied randomly around the hardness values of the workpiece material prior to
machining. Machining with PCD tools generally led to softening of machined surfaces. Increase in
cutting speed generally led to increased hardness when machining with the larger grain size PCD (T5)
tool using conventional coolant flow and with coolant pressures up to 11 MPa. No evidence of plastic
deformation was observed on the machined surfaces and the surface integrity of the finish machined
surfaces is generally in agreement with Rolls–Royce CME 5043 specification.
Keywords: Titanium alloy, High Coolant Pressure, Various cutting tools, Tool life, Surface integrity
viii
DA SILVA, R.B. Desempenho de diferentes Materiais de Ferramentas de Corte no Torneamento de Acabamento
da liga de titânio Ti-6Al-4V com a Tecnologia de Aplicação de Fluido de Corte à Alta Pressão, 2006, 299 f. Tese
de Doutorado, Universidade Federal de Uberlândia, Uberlândia.
RESUMO
Este estudo visa avaliar a usinabilidade da liga de titânio Ti-6Al-4V utilizando várias classes de diferentes
materiais de ferramentas de corte tais como metal duro sem revestimento (insertos T1 e T3) e com revestimento
(insertos T2 e T4), PCD – insertos: T5 e T6, CBN – insertos: T7,T8 e T9, cerâmicas Whiskers (inserto T10), e
nano-cerâmicas à base de alumina (inserto T11) e à base de nitreto de silício (inserto T12) em diferentes
atmosferas de usinagem (fluido de corte aplicado a altas pressões (HPC) de 7 MPa; 11 MPa and 20,3 MPa,
argônio e aplicação de fluido de corte convencional) e em elevadas condições de corte típicas de acabamento
(velocidade de corte de 100 m min
-1
a 500 m min
-1
, com avanço de 0,15 mm volta
-1
e profundidade de corte
de 0,5 mm constantes). Foram monitorados a vida das ferramentas bem como os mecanismos e tipos de desgaste,
as forças de usinagem, a integridade superficial, a rugosidade das superfícies usinadas, a circularidade e os tipos
e classes de cavacos produzidos. Os resultados foram utilizados para avaliar a eficiência das diferentes
ferramentas de corte e atmosferas de usinagem empregadas na usinagem da liga Ti-6Al-4V. Os resultados
mostraram que as ferramentas de PCD e metal duro tiveram o melhor desempenho, em termos de vida de
ferramenta, que as demais ferramentas testadas. Em geral, as ferramentas com tamanho de grãos maior, metal
duro (T1 e T4) e PCD (T5), apresentaram o melhor desempenho, em termos baixa taxa de desgaste e,
consequentemente, vida mais longa, comparada com as ferramentas com tamanho de grãos menores (classes
T2,T3 e T6). A utilização da técnica HPC mostrou ser eficiente na usinagem da liga Ti-6Al-4V, em termos de
aumento de vida da ferramenta e, consequentemente, de aumento de produtividade, em relação à técnica de
aplicação de fluido de corte convencional e com utilização de argônio nas condições investigadas. Em geral, a
vida das ferramentas aumentaram com o aumento da pressão de aplicação de fluido de corte devido à sua
capacidade de reduzir a área de contato cavaco-ferramenta e de quebrar o cavaco mais eficientemente e,
portanto, propiciando uma melhor condição de lubrificação na interface cavaco-ferramenta com conseqüente
redução de atrito. A utilização do argônio na usinagem com as ferramentas T1, T4 e T10 nas condições
investigadas apenas evitou com que o centelhamento e ignição do titânio ocorresse, além de não propiciar
aumento de vida da ferramenta, provavelmente devido à supressão das características de refrigeração e
lubrificação que o argônio tem. As ferramentas de PCD apresentaram uma vida cerca de 8 vezes maior que as
ferramentas de metal duro quando empregadas com aplicação de fluido de corte convencional. Todas as classes
de ferramentas de CBN, em geral, apresentaram baixo desempenho em termos de vida de ferramenta devido ao
acelerado desgaste na ponta da ferramenta e, em certos casos, lascamentos da aresta de corte que estão
associados com a relativa alta taxa de difusão que ocorre durante a usinagem com titânio, que tende a diminuir a
forças de ligações entre os átomos do substrato. Todas as ferramentas de cerâmicas testadas não demonstraram
desempenho satisfatório em termos de desgaste e de vida ferramenta durante a usinagem da liga Ti-6Al-4V por
causa da ocorrência de desgaste abrasivo e de lascamento da aresta de corte, como também da produção de
superfícies usinadas com pobre acabamento superficial. O desgaste de ponta foi o tipo de desgaste predominante
durante a usinagem com as ferramentas de metal duro, PCD e CBN (T7) devido à redução da área de contato
cavaco-ferramenta e, consequentemente, ao aumento das tensões atuantes e aumento da temperatura na ponta da
ferramenta. Já o desgaste de entalhe e lascamento ocorreram durante a usinagem com as ferramentas de CBN
(T8 and T9) e com cerâmicas convencionais. O desgaste de entalhe também ocorreu de forma mais acentuada
nas ferramentas de nano-cerâmicas, o que levou à falha catastrófica de tais ferramentas quando empregadas em
velocidades de corte superiores a 110 m min
-1
. A usinagem com ferramentas de PCD geraram baixas forças de
corte em relação às ferramentas de metal duro. Os valores de rugosidade superficial produzidos com as
ferramentas de metal duro, PCD e CBN em geral ficaram abaixo do valor estipulado para critério de rejeição
para torneamento de acabamento de 1.6 µm, enquanto que todas as ferramentas de cerâmicas produziram valores
de rugosidade acima de 2 µm. A análise metalográfica das superfícies usinadas permitiu identificar pequenas
marcas que não comprometeram as superfícies produzidas. A usinagem com ferramentas de metal duro produziu
valores de dureza que variam aleatoriamente dentro dos limites inferior e superior de dureza da peça medidos
antes da usinagem. Nenhuma evidência de deformação plástica nas superficies de titânio usinadas com todas as
ferramentas e condições testadas. Em geral, a integridade superficial das superficies usinadas atendem à norma
Rolls–Royce CME 5043.
Palavras-chave: Liga de titânio, Fluido de corte à alta pressão, Várias ferramentas de corte, Vida de ferramenta,
Integridade superficial.
ix
TABLE OF CONTENTS
FIGURES…………………………………………………………………………... xv
TABLES ……………………………………………………………………….…... xxvi
LIST OF SYMBOLS ...……………………………………………………….…... xxvii
Chapter I INTRODUCTION …………………………………………………... 1
1.1 Aims of the Thesis 4
Chapter II LITERATURE SURVEY 6
2.1 Historical Background of Machining 6
2.2 Overview of Aerospace Alloys 8
2.2.1 Aero-Engine Alloys 9
2.3 Superalloys 14
2.3.1 Titanium Superalloys in the Aerospace Industry 15
2.4 Machining Operations 23
2.4.1 Terminology used in Metal Cutting 24
2.4.2 Nomenclature of Cutting Tools 28
2.5 Chip Formation Process 30
2.6 Classes of Chips 33
2.7 Forces in Metal Cutting 36
2.8 Stress and Strain Distribution in Machining 39
2.8.1 Stress Distribution 39
2.8.2 Strain Distribution 40
2.9 Heat Generation During Machining Operation 42
2.9.1 Effect of Cutting Parameters On Temperature Generated
During Machining
43
2.9.2 Heat Generation and Cutting Temperature when
Machining Titanium Alloy
45
2.10 Tool Failure Modes 47
2.11 Tool Wear Mechanisms 49
2.12 Titanium Machinability 51
2.13 Tool Materials for Machining Titanium Alloys 53
x
2.13.1 Tool Materials Requirements 55
2.13.2 High Speed Steel (HSS) Tools 56
2.13.3 Cemented Carbide Tools 59
2.13.3.1 Uncoated Carbide Tools 62
2.13.3.2 Coated Carbide Tools 67
2.13.4 Ultrahard (Superabrasive) Tool Materials 72
2.13.4.1 Polycrystalline Diamond (PCD) Tools 73
2.13.4.2 Cubic Boron Nitride (CBN) Tools 76
2.13.5 Ceramic Tools 80
2.13.5.1 Pure Oxide Ceramics 81
2.13.5.2 Mixed Oxide Ceramics 81
2.13.5.3 Whisker Reinforced Alumina Ceramics 82
2.13.5.4 Silicon Nitride-base Ceramics 82
2.13.5.5 Nano-grain Ceramics 84
2.14 Cutting Fluids 86
2.14.1 Classification of Cutting Fluids 88
2.14.2 Directions of Application of Cutting Fluids 92
2.15 Cutting Environments and Techniques Employed when Machining
Titanium Alloys
94
2.15.1 Dry Machining 94
2.15.2 Conventional Coolant Supply 99
2.15.3 High Pressure and Ultra Pressure Coolant Supplies 100
2.15.4 Minimum Quantity of Lubrication (MQL) 106
2.15.5 Cryogenic Machining 108
2.15.6 Other Atmospheres 111
2.15.7 Ledge Cutting Tools 114
2.15.8 Rotary Tools 115
2.15.9 Ramping Technique 117
2.15.10 Hot Machining / Hybrid Machining 118
2.16 Surface Integrity 121
2.16.1 Surface Finish and Texture 121
2.16.2 Subsurface Changes 124
xi
Chapter III EXPERIMENTAL PROCEDURE 126
3.1 Introduction 126
3.2 Work Material 127
3.3 Machine Tool 127
3.4 Cutting Fluid 128
3.5 High Pressure Unit 128
3.6 Argon Delivery System 130
3.7 Tool Material and Machining Procedure 131
3.8 Cutting Conditions 135
3.9 Tool Life Criteria 137
3.10 Tool Wear Measurement 137
3.11 Component Force Measurement 138
3.12 Surface Roughness Measurement 139
3.13 Runout Measurement 141
3.14 Tool and Workpiece Specimen Preparation 141
3.15 Microhardness Measurements Below the Machined Surface 143
Chapter IV EXPERIMENTAL RESULTS 145
4.1 Benchmark trials - Machining of Ti-6Al-4V alloy with Uncoated
Carbide (883 grade) inserts
145
4.2 Machining of Ti-6Al-4V alloy with various carbide tool grades
(uncoated and coated tools) under various machining environments
146
4.2.1 Tool life 146
4.2.2 Tool wear when machining Ti-6Al-4V alloy with various
carbide insert grades
149
4.2.3 Component forces when machining with various carbide
insert grades
159
4.2.4 Surfaces roughness and runout values when machining
with various carbide insert grades
161
4.2.5 Surfaces generated after machining with various carbide
insert grades
163
4.2.6 Surface hardness after machining with various carbide
tool grades
167
4.2.7 Subsurface micrographs after machining Ti-6Al-4V alloy
with various carbide insert grades
175
4.2.8 Chips shapes 180
4.3
Machining of Ti-6Al-4V alloy with different grades of PCD tools
under various coolant supply pressures
182
4.3.1 Tool life 182
xii
4.3.2 Tool wear when machining Ti-6Al-4V alloy with
different grades of PCD tools
184
4.3.3 Component forces when machining with different grades
of PCD tools
190
4.3.4 Surfaces roughness and roundness values when
machining with different grades of PCD tools
192
4.3.5 Surface alteration after machining with different grades
of PCD tools
194
4.3.6 Surface hardness after machining with different grades of
PCD tools
197
4.3.7 Subsurface alteration after machining Ti-6Al-4V alloy
with different grades of PCD tools
200
4.3.8 Chips shapes 203
4.4 Machining of Ti-6Al-4V alloy with different grades of CBN tools
under various coolant supply pressures
204
4.4.1 Tool life 204
4.4.2 Tool wear when machining Ti-6Al-4V alloy with
different grades of CBN tools
206
4.4.3 Component forces 211
4.4.4 Surfaces roughness and runout values 212
4.4.5 Surface hardness and subsurface alteration 214
4.4.6 Chips shapes 216
4.5 Machining of Ti-6Al-4V alloy with whisker reinforced ceramic
cutting tools under various machining environments
217
4.5.1 Wear rate and tool life 217
4.5.2 Component forces 222
4.5.3 Surfaces roughness 223
4.5.4 Surface hardness and subsurface alterations 224
4.5.5 Chips shapes 226
4.6 Machining of Ti-6Al-4V alloy with Nano-ceramic cutting tools 227
4.6.1 Wear rate and tool life 227
4.6.2 Component forces 230
4.6.3 Surfaces roughness and runout values 230
4.6.4 Chips shapes 232
Chapter V DISCUSSIONS 233
5.1 Introduction 233
5.2 Tool performance when machining Ti-6Al-4V alloy with different
grades of carbide, PCD, CBN and ceramic tools
233
xiii
5.2.1 Carbides tools 233
5.2.2 PCD tools 237
5.2.3 CBN tools 239
5.2.4 Micron-grain ceramic tools 240
5.2.5 Nanoceramic tools 241
5.3 Tool failure modes and wear mechanisms when machining Ti-6Al-
4V alloy with different grades of carbide, PCD, CBN and ceramic
tools
242
5.3.1 Carbide tools 242
5.3.2 PCD tools 246
5.3.3 CBN tools 248
5.3.4 Micron-grain ceramic tools 251
5.3.5 Nano-grain ceramic tools 253
5.4 Components forces when machining Ti-6Al-4V alloy with different
grades of carbide, PCD, CBN and ceramic tools
254
5.5 Surfaces roughness and runout values when machining Ti-6Al-4V
alloy with different grades of carbide, PCD, CBN and ceramic tools
258
5.6 Surface generated of Ti-6Al-4V after machining with carbide and
PCD tools
260
5.7 Surface hardness after machining Ti-6Al-4V alloy with different
grades of carbide, PCD, CBN and ceramic tools
261
5.8 Subsurface micrographs after machining Ti-6Al-4V alloy with
different grades of carbide, PCD, CBN and ceramic tools
265
5.9 Chips shapes 266
Chapter VI CONCLUSIONS 269
Chapter VII RECOMMENDATIONS FOR FURTHER WORK 272
Chapter VIII REFERENCES 274
APPENDIX 298
LIST OF PUBLICATIONS FROM THIS STUDY 298
Refereed Journals 298
Refereed Conferences 299
xiv
FIGURES
Figure 2.1 - (a) A typical jet engine and its main parts (Pratt and Whitney F100 jet
engine) (BENSON, 2002); (b) typical jet engine (Trent 700) manufactured by Rolls-
Royce plc (ROLLS-ROYCE PLC, 2003)
10
Figure 2.2 - Trends in turbine inlet temperature in areo-engines (OHNABE et al., 1999) 11
Figure 2.3 - Improvements in aero-engine performance (BENSON, 2002) 11
Figure 2.4 - Maximum service temperature of various materials (LOVATT;
SHERCLIFF, 2002)
12
Figure 2.5 - Trend of materials usage in aero-engines (MILLER, 1996) 12
Figure 2.6 - Typical applications of titanium: (a) modular femoral components
(prostheses manufactured in titanium-base, Ti-6Al-4V, alloy, (b) valves and (c) screw
(TIG, 2002)
20
Figure 2.7 - Phases of titanium product life cycle in the U.S. (ASM HANDBOOK,
1998)
20
Figure 2.8 - Metal cutting diagram (WATERS, 2000)
23
Figure 2.9 - Basic machining operation and important parameters (KALPAKJIAN;
SCHMID, 2000)
25
Figure 2.10 - Schematic illustration of typical single-point cutting tool with the tool
angles (KALPAKJIAN; SCHMID, 2000)
26
Figure 2.11 - Workpiece-tool-machine system for turning operation
26
Figure 2.12 - Form-milling operation with gangs of side and face milling cutters (AB
SANDVIK COROMANT, 1994)
27
Figure 2.13 - Cutting tool planes: (a) “tool-in-hand” planes and (b) “tool-in-use” planes
(BOOTHROYD; KNIGHT, 1989)
29
Figure 2.14 - Tool angles for a single-point tool according to the ISO: tool cutting edge
angle (k
r
), tool minor cutting edge angle (k´
r
), tool included angle (
ε
r
), tool cutting edge
inclination angle (
λ
s
), tool normal rake angle (γ
n
), tool normal clearance angle (
α
n
) and
tool normal wedge angle (
β
n
) (BOOTHROYD; KNIGHT, 1989)
30
Figure 2.15 - Metal cutting diagram - the chip formation (TRENT; WRIGHT, 2000)
32
Figure 2.16 - Metal cutting diagram illustrating the primary and secondary shear zones
(THE METALS HANDBOOK, 1989)
32
Figure 2.17 - Classes of chips: (a) Continuous chip, (b) Continuous chip with BUE,
(c) Discontinuous chip, (d) Serrated chips (TRENT; WRIGHT (2000), MACHADO;
WALLBANK (1990), KALPAKJIAN; SCHMID (2000))
35
Figure 2.18 - Cutting forces a) Three components forces acting on the cutting tool (DE
GARMO; BLACK; KOHSER, 1999) and b) Merchant´s circle (TRENT; WRIGHT,
2000)
38
Figure 2.19 - The Zorev´s model of stress distribution on the rake face of a cutting tool
in orthogonal cutting where σ
fmax
= maximum normal stress, σ
f
= normal stress, τ
f
= shear
stress, τ
st
= shear strength of chip material in the sticking region (BOOTHROYD;
KNIGHT, 1989)
40
Figure 2.20 - The shear strain in the shear plane (SHAW, 1984)
41
Figure 2.21 - Zones of heat generation during machining: (a) schematic diagram,
(b) isothermal lines for dry orthogonal cutting of free machining steel with carbide tool
(α = 20º) obtained from a finite element technique, at a cutting speed of 155.4 m min
-1
43
xv
and a feed rate of 0.274 mm rev
-1
[adapted from (SHAW, 1984)]
Figure 2.22 - Distribution of thermal load when machining titanium-base, Ti-6Al-4V
and steel Ck 45 [adapted from (DEARNLEY; GREARSON, 1986)
46
Figure 2.23 - Influence of cutting speed on the cutting temperature when machining
titanium and its alloys [adapted from (MOTONISHI et al., 1987)]
46
Figure 2.24 - Regions of wear on a cutting tool (DEARNLEY; TRENT, 1985) 47
Figure 2.25 - The main wear mechanisms on a cutting tool [adapted from (TRENT;
WRIGHT, 2000)]
50
Figure 2.26 - Influence of temperature on hot hardness of some tool materials
(ALMOND, 1981)
56
Figure 2.27 - Flow stress measured at 0.6% strain during three point bending tests at a
constant strain rate in WC-11wt.%Co (MARI; GONSETH, 1993)
60
Figure 2.28 - Tool live when turning Ti-6242 alloy with mixed uncoated carbide tools
with different grain sizes of substrates: 0.68 µm (890 grade) and 1.0 µm (883 grade)
(JAWAID; CHE-HARON; ABDULLAH, 1999)
61
Figure 2.29 - Average crater wear rates of various tool materials in turning of Ti-6Al-4V
alloy at a cutting speed of 61 m min
-1
for 10 minutes (HARTUNG; KRAMER, 1982)
64
Figure 2.30 - SEM micrograph of exposed mixed cemented carbide substrate after
fracture of the welded junction (NABHANI, 2001b)
65
Figure 2.31 - Flank face of a worn uncoated straight carbide tool showing abrasion by
carbide grains after turning titanium base, Ti-6242, alloy under dry condition (JAWAID;
CHE-HARON; ABDULLAH, 1999)
65
Figure 2.32 - Evidence of adhesion of the chips on the nose of a mixed ((Ta,Nb)C)
uncoated straight carbide tool after machining Ti-6Al-4V alloy with conventional
coolant supply at a speed of 100 m min
-1
, a feed rate of 0.15 mm rev
-1
and a depth of cut
of 0.5 mm (EZUGWU et al., 2005)
67
Figure 2.33 - Flank wear curves when machining Ti-6Al-4V alloy with coated (CrN and
TiCN) and a straight uncoated carbide tools (TURLEY, 1981)
71
Figure 2.34 - Worn surface of a multilayer (TiC/TiCN/TiN) coated mixed cemented
carbide tool showing remains of adherent metal layer (a) and enlarged view of the crater
wear showing smooth ridges with fine scoring in direction of chip flow (b) after
machining titanium base, Ti-5Al-4Mo-(2-2.5)Sn-(6-7)Si alloy (NABHANI, 2001b)
71
Figure 2.35 - Coating delamination of PVD coated (TiN) carbide tool, grinding marks
and adhered material observed after 10 s (a) and adhesion of work material onto the
flank face, plastic deformation and cracks at the cutting edge after 20 s; (b) after face
milling Ti-6Al-4V alloy at cutting speeds of 100 m min
-1
and 50 m min
-1
and feed rates
of 0.15 mm per tooth and 0.1 mm per tooth, respectively (JAWAID; SHARIF;
KOKSAL, 2000)
72
Figure 2.36 - The performance of various grades of PCD tools when milling ceramic
impregnated surface of a flooring board (HPL) (COOK; BOSSOM, 2000)
75
Figure 2.37 - Formation of strongly adherent layer on the rake face of a PCD tool after
machining titanium base, Ti-5Al-4Mo-2Sn-6Si alloy under dry condition (NABHANI,
2001b)
76
Figure 2.38 - (a) Section through ‘quick-stop’ specimen showing part of CBN tool
adhering to underside of chip (100x), (b) close-up view of Fig. 2.38(a) (200x)
(NABHANI, 2001a)
78
Figure 2.39 - (a) A typical scanning electron micrographs of worn-out edges: (a) cutting
temperature of 734ºC, (b) cutting temperature of 900ºC (ZOYA; KRISHNAMURTHY,
2000)
80
xvi
Figure 2.40 - Variation in uniform flank wear with cutting time for the turning of Ti-
6Al-4V (hardness, 36 HRC), showing reduced tool wear with the new geometry (cutting
speed, 122 m min
-1
; feed rate, 0.23 mm rev
-1
unless otherwise indicated; depth of cut,
1.52 mm; tool SNG432 (SCEA, 15º): curve A, SIALON (Kyon 2000) with clearance
angles of 17º (localised wear, 0.889 mm; edge fracture; crater) and 5º (localized wear,
1.321 mm; fracture; crater); curve B, SIALON (Kyon 2000) with a clearance angle of 5º
and a feed rate of 0.127 mm rev
-1
; curves C and D, cemented carbide (Carboloy grade
999) with clearance angles of 5º and 17º, respectively (KOMANDURI; REED JR, 1983)
84
Figure 2.41 - TEM micrograph of HIPed (Hot Isostatic Pressing) nanophase SiC sample
with a density of 97% TD (Theoretical density) (VAβEN; STÖVER, 1999)
85
Figure 2.42 - Schematic illustration of the possible directions of application of cutting
fluids
93
Figure 2.43 - Schematic illustration of a tool holder used for machining with high
pressure coolant supply (SECO TOOLS, 2002b)
100
Figure 2.44 - Pressure distribution from the jet momentum action on the chip. (a)
Cutting in tube with single straight edge; (b) pressure distribution (2D) at longitudinal
turning (DAHLMAN, 2000).
102
Figure 2.45 - Schematic illustration of nozzle orientation for localized LN2 delivery
(HONG; DING; JEONG, 2001)
109
Figure 2.46 - A schematic representation of the cryogenic cooling concepts
(MAZURKIEWICZ; KUBALA; CHOW, 1989)
110
Figure 2.47 -The tool assembly, nozzles and LN2 flowing out of the nozzle
(MAZURKIEWICZ; KUBALA; CHOW, 1989)
110
Figure 2.48 - Ledge tool (after KOMANDURY; LEE, 1984))
115
Figure 2.49 - Schematic representation of principle of rotary cutting action (WANG;
EZUGWU; GUPTA, 1998)
116
Figure 2.50 - Schematic representation of a hot machining technique design (ÖZLER;
ÍNAN; ÖZEL, 2001)
120
Figure 2.51 - Standard terminology and symbols of the elements of surface texture (µin)
(KALPAKJIAN; SCHMID, 2000)
122
Figure 2.52 - Schematic illustration of the determination of some amplitude parameters
of surface texture (SHOUCKRY; 1982)
123
Figure 2.53 - Form tolerances for machined surfaces in turning operations: (a)
Roundness, (b) Cylindricity (DE GARMO; BLACK; KOHSER, 1999)
126
Figure 2.54 - Production and cost curves versus cutting speed (GORCZYCA, 1987)
128
Figure 3.1 - Colchester Electronic MASTIFF CNC lathe
128
Figure 3.2 - The high pressure pumping coolant system - Chipblaster (CV26-3000)
129
Figure 3.3 - Special tool holder and a cutting fluid jet-pressure of 7 MPa supply.
130
Figure 3.4 - Argon gas delivery system: (a) cylinder and (b) close-up view of the valve
and the hose.
131
Figure 3.5 - Cutting tools used in the machining trials: uncoated carbides: T1 (883
grade), T2 (890 grade), coated carbides T3 ( CP 200), T4 (CP 250 grade); PCD: T5 (20
grade with grain size of 10 µm), T6 (20 grade with grain size < 10 µm); CBN: T7 (10
grade), T8 (300 grade), T9 (300-P grade); silicon carbide (SiC
w
) whisker reinforced
alumina ceramic inserts (WG300): T10 (rhomboid shaped) and T11 (square shaped);
nano-grain ceramic inserts: T12 (Al
2
O
3
grade) and T13 (Si
3
N
4
grade).
132
xvii
Figure 3.6 - Tool holders used in the machining trials: (a) designation PCLNR2525-M12
used for carbide tools (T1,T2,T3,T4); (b) designation SCLCR2525-M12 used for PCD
tools (T5,T6); (c) designation DCLNR2525-M12 used for CBN and ceramic tools
(T7,T8, T9 and T10); (d) designation MSLNR-252512 used for square tools: micron-
grain and nano-grain size ceramics (T11,T12,T13).
134
Figure 3.7 - Mitutoyo tool maker´s microscope.
138
Figure 3.8 - (a) Kistler dynamometer for capturing forces generated during machining
and (b) Oscilloscope with charge amplifier.
139
Figure 3.9 - (a) Surtronic-10 portable stylus type used for surface roughness
measurement; (b) dial indicator Shockproof – BATY used for roundness measurement.
140
Figure 3.10 - (a) Hitachi (S530) Scanning Electron Microscope; (b) Nicon Metallurgical
Optical Microscope (OPTIPHOT-100) with computerised image system.
142
Figure 3.11 - (a) Buehler Automatic Mounting Press (Simplimet 2000); (b) Automatic
Grinding/Polishing Equipment (Metaserv 2000).
143
Figure 3.12 - Mitutoyo (MVK – VL) Vickers micro-hardness tester machine.
144
Figure 4.1 - Average flank wear of uncoated carbide (T1) insert at various cutting speeds
with conventional coolant supply during 15 minutes machining time (benchmark trials)
146
Figure 4.2 - Figure 4.2 - Tool life recorded when machining Ti-6Al-4V alloy with
different cemented carbide insert grades with conventional coolant flow (CCF), high
coolant pressures of 7 MPa, 11 MPa and 20.3 MPa and in argon enriched environment
at various speed conditions.
147
Figure 4.3 - Nose wear rate curves of different cemented carbide insert grades when
machining Ti-6Al-4V alloy with conventional coolant flow (CCF), high coolant
pressures of 7 MPa, 11 MPa and 20.3 MPa and in argon enriched environment, at a feed
rate of 0.15 mm rev
-1
and a depth of cut of 0.5 mm.
150
Figure 4.4 - Nose wear curves when finish machining with cemented carbide (T1 and
T4) inserts at a cutting speed of 110 m min
-1.
151
Figure 4.5 - Nose wear curves when finish machining with cemented carbide (T2 and
T3) inserts at a cutting speed of 110 m min
-1
.
151
Figure 4.6 - Worn T1 insert after machining Ti-6Al-4V alloy with conventional coolant
supply at a speed of (a) 100 m min
-1
and (b) 130 m min
-1.
152
Figure 4.7 - Wear generated at the cutting edge of uncoated carbide T1 insert after
machining Ti-6Al-4V alloy with a coolant pressure of 7MPa at a speed of 110 m min
-1.
153
Figure 4.8 -Wear generated at the cutting edge of uncoated carbide T1 insert after
machining Ti-6Al-4V alloy with a coolant pressure of 11MPa at a speed of 120 m min
-1
.
153
Figure 4.9 - Worn cutting edge of uncoated carbide T1 insert after machining Ti-6Al-4V
alloy with a coolant pressure of 20.3 MPa at a speed of 130 m min
-1
.
154
Figure 4.10 - Worn cutting edge of uncoated carbide T1 insert after machining Ti-6Al-
4V alloy in argon enriched environment at a speed of 130 m min
-1
.
154
Figure 4.11 - Flank and nose wears at the cutting edge of uncoated carbide T2 insert
grade after machining Ti-6Al-4V alloy with conventional coolant supply at a speed of
130 m min
-1
, a feed rate of 0.15 mm rev
-1
and a depth of cut of 0.5 mm.
155
Figure 4.12 - Wear generated at the cutting edge of uncoated carbide T2 insert after
machining Ti-6Al-4V alloy with a coolant pressure of 11 MPa at a speed of
(a) 110 m min
-1
and (b) 130 m min
-1.
155
Figure 4.13 - Wear generated at the cutting edge of uncoated carbide T2 insert after
machining Ti-6Al-4V alloy with a coolant pressure of 20.3 MPa at a speed of
130 m min
-1
.
156
xviii
Figure 4.14 - Worn cutting edge of T3 coated carbide insert when machining with
conventional coolant supply at a speed of (a) 110 m min-1 and (b) 130 m min
-1
.
156
Figure 4.15 - Flank and nose wears a the cutting edge of T3 coated carbide insert after
machining Ti-6Al-4V alloy with a coolant pressure of 11 MPa at a speed of
110 m min
-1
.
157
Figure 4.16 - Flank and nose wears at the cutting edge of T3 coated carbide insert after
machining Ti-6Al-4V alloy with a coolant pressure of 20.3 MPa at a speed of
(a) 110 m min
-1
and (b) 130 m min
-1
.
157
Figure 4.17 - Worn cutting edge of T4 coated carbide insert after machining Ti-6Al-4V
alloy with conventional coolant supply at a speed of 130 m min
-1
.
158
Figure 4.18 - Adhesion of work material on a worn T4 coated carbide insert after
machining Ti-6Al-4V alloy with a coolant pressure of 11 MPa at a speed of
110 m min
-1
.
158
Figure 4.19 - Nose wear at the cutting edge of T4 coated carbide insert after machining
Ti-6Al-4V alloy with a coolant pressure of 20.3 MPa at a speed of (a) 110 m min
-1
and
(b) 120 m min
-1
.
159
Figure 4.20 - Wear at the cutting edge of T4 coated carbide insert after machining Ti-
6Al-4V alloy in argon enriched environment at a speed of (a) 100 m min
-1
and
(b) 120 m min
-1
.
159
Figure 4.21 - Cutting forces (Fc) recorded at the beginning of cut when machining Ti-
6Al-4V alloy with different cemented carbide grades with various cutting conditions.
160
Figure 4.22 - Feed forces (F
f
) recorded at the beginning of cut when machining Ti-6Al-
4V alloy with different cemented carbide grades under various cutting conditions.
161
Figure 4.23 - Surface roughness values recorded at the beginning of cut when machining
Ti-6Al-4V alloy with different cemented carbide grades under various cutting
conditions.
162
Figure 4.24 - Roundness variation recorded at the end of cut when machining Ti-6Al-4V
alloy with different cemented carbide grades under various cutting conditions.
163
Figure 4.25 - Surfaces generated after machining with uncoated carbide T1 tool with
conventional coolant supply at cutting speeds of (a) 110 m min
-1
and (b) 130 m min
-1
.
163
Figure 4.26 - Surfaces generated after machining with uncoated carbide T1 tool with a
coolant pressure of 7 MPa at cutting speeds of (a) 100 m min
-1
and (b) 130 m min
-1
.
164
Figure 4.27 - Surfaces generated after machining with uncoated carbide T1 tool with a
coolant pressure of 11 MPa at cutting speeds of (a) 110 m min
-1
and (b) 120 m min
-1
.
164
Figure 4.28 - Surfaces generated after machining with uncoated carbide T1 tool with a
coolant pressure of 20.3 MPa at cutting speeds of (a) 120 m min
-1
and (b) 130 m min
-1
.
164
Figure 4.29 - Surfaces generated after machining with uncoated carbide T1 tool in an
argon enriched environment at cutting speeds of (a) 110 m min
-1
and (b) 120 m min
-1
.
165
Figure 4.30 - Surfaces generated after machining with uncoated carbide T2 tool with
coolant pressures of (a) 11 MPa and (b) 20.3 MPa at a cutting speedsof (a) 110 m min
-1
.
165
Figure 4.31 - Surfaces generated after machining with coated carbide T3 tool with
coolant pressures of (a) 11 MPa and (b) 20.3 MPa at a cutting speed of 110 m min
-1
.
166
Figure 4.32 - Surfaces generated after machining with coated carbide T4 tool with (a)
conventional coolant supply, (b) in argon enriched environment, (c) coolant pressure of
11 MPa and (d) 20.3 MPa at a cutting speed of 120 m min
-1
.
167
Figure 4.33 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide
(T1) insert grade with conventional coolant supply.
169
Figure 4.34 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide
(T1) insert grade with 7 MPa coolant pressure.
169
xix
Figure 4.35 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide
(T1) insert grade with 11MPa coolant pressure.
170
Figure 4.36 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide
(T1) insert grade with 20.3 MPa coolant pressure.
170
Figure 4.37 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide
(T1) insert grade in argon-enriched environment.
171
Figure 4.38 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide
(T2) insert grade with various cutting conditions.
172
Figure 4.39 - Hardness variation after machining Ti-6Al-4V alloy with coated carbide
(T3) insert grade with various cutting conditions.
172
Figure 4.40 - Hardness variation after machining Ti-6Al-4V alloy with coated carbide
(T4) insert grade with conventional coolant.
173
Figure 4.41 - Hardness variation after machining Ti-6Al-4V alloy with coated carbide
(T4) insert grade with 11MPa coolant pressure.
174
Figure 4.42 - Hardness variation after machining Ti-6Al-4V alloy with coated carbide
(T4) insert grade with 20.3 MPa coolant pressure.
174
Figure 4.43 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide
T1 inserts with conventional coolant supply at cutting speeds of (a) 110 m min
-1
and
(b) 130 m min
-1
.
175
Figure 4.44 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide
T1 inserts with a coolant pressure of 7 MPa at cutting speeds of (a) 100 m min
-1
and
(b) 120 m min
-1
.
175
Figure 4.45 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide
T1 inserts with a coolant pressure of 11 MPa at cutting speeds of (a) 110 m min
-1
and (b)
120 m min
-1
.
176
Figure 4.46 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide
T1 insert with a coolant pressure of 20.3 MPa at cutting speeds of (a) 120 m min
-1
and
(b) 130 m min
-1
.
176
Figure 4.47 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide
T1 inserts in an argon enriched environment at cutting speeds of (a) 110 m min
-1
and
(b) 120 m min
-1
.
176
Figure 4.48 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide T2 inserts with
conventional coolant supply at a cutting speed of (a) 110 m min
-1
and (b) 130 m min
-1
.
177
Figure 4.49 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide
T2 inserts with a coolant pressure of 11 MPa at cutting speeds of (a) 110 m min
-1
and
(b) 130 m min
-1
.
177
Figure 4.50 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide
T2 tools with a coolant pressure of 20.3 MPa at cutting speeds of (a) 110 m min
-1
and
(b) 130 m min
-1
.
177
Figure 4.51 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T3 inserts with
conventional coolant supply at a cutting speed of (a) 110 m min
-1
and (b) 130 m min
-1
.
178
Figure 4.52 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T3
tools with a coolant pressure of 11 MPa at cutting speeds of (a) 110 m min
-1
and
(b) 130 m min
-1
.
178
Figure 4.53 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T3
tools with a coolant pressure of 20.3 MPa at cutting speeds of (a) 110 m min
-1
and
(b) 130 m min
-1
.
178
Figure 4.54 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T4
tools with conventional coolant supply at cutting speeds of (a) 100 m min
-1
and
(b) 120 m min
-1
.
179
xx
Figure 4.55 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T4
tools with a coolant pressure of 11 MPa at cutting speeds of (a) 120 m min
-1
and
(b) 130 m min
-1
.
179
Figure 4.56 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T4
tools with a coolant pressure of 20.3 MPa at cutting speeds of (a) 110 m min
-1
and
(b) 130 m min
-1
.
179
Figure 4.57 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T4
tools in an argon enriched environment at cutting speeds of (a) 120 m min
-1
and
(b) 130 m min
-1
.
180
Figure 4.58 - Chips generated when machining Ti-6Al-4V alloy with different carbide
tool grades under various cutting conditions: (a) continuous tubular chip; (b), (f), (h) and
(k) continuous and snarled chips, (c), (g) and (i) partially segmented chips, (d), (e) and
(l) segmented C-shaped chips.
181
Figure 4.59 - Tool life recorded when machining Ti-6Al-4V alloy with PCD-STD (T5)
and PCD MM (T6) tool grades with conventional coolant flow (CCF) and high coolant
pressures of 7 MPa, 11 MPa and 20.3 MPa at various cutting speed conditions.
183
Figure 4.60 - Nose wear rate when machining Ti-6Al-4V alloy with PCD inserts with
conventional coolant flow and high coolant pressures of 7 MPa, 11 MPa and 20.3 MPa
at various cutting speed conditions.
185
Figure 4.61 - Nose wear when finish machining with PCD-STD (T5) and PCD-MM
insert grades (T6) at a cutting speed of 175 m min
-1
.
186
Figure 4.62 - Wear observed on T5 insert after machining Ti-6Al-4V alloy with
conventional coolant supply at a speed of (a) 140 m min
-1
and (b) 200 m min
-1
.
187
Figure 4.63 - Worn T5 insert after machining Ti-6Al-4V alloy with a 7MPa coolant
pressure and at a speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
187
Figure 4.64 - Worn T5 insert after machining Ti-6Al-4V alloy with 11 MPa coolant
pressure at a speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
188
Figure 4.65 - Wear observed on a T5 insert after machining Ti-6Al-4V alloy with 20.3
MPa coolant pressure at a speed of (a) 200 m min
-1
and (b) 250 m min
-1
.
188
Figure 4.66 - Wear observed on a T6 insert after machining Ti-6Al-4V alloy with
conventional coolant supply at a speed of 175 m min
-1
.
189
Figure 4.67 - Worn T6 insert after machining Ti-6Al-4V alloy with 11MPa coolant
pressure at a speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
189
Figure 4.68 - Worn T6 insert after machining Ti-6Al-4V alloy with 20.3 MPa coolant
pressure at a speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
190
Figure 4.69 - Cutting forces (Fc) recorded at the beginning of cut when machining Ti-
6Al-4V alloy with PCD-STD (T5) and PCD-MM insert grades (T6) at various cutting
conditions.
191
Figure 4.70 - Feed forces (Ff) recorded at the beginning of cut when machining Ti-6Al-
4V alloy with PCD-STD (T5) and PCD-MM insert grades (T6) at various cutting
conditions.
191
Figure 4.71 - Surface roughness values recorded at the beginning of cut when machining
Ti-6Al-4V alloy with T5 and T6 inserts at various cutting conditions.
193
Figure 4.72 - Roundness values recorded at the end of cut after machining Ti-6Al-4V
alloy with T5 and T6 insert grades at various cutting conditions.
193
Figure 4.73 - Surfaces generated after machining with PCD (T5) inserts with
conventional coolant supply at a cutting speed of (a) 175 m min
-1
and (b) 200 m min
-1
.
195
Figure 4.74 - Surfaces generated after machining with PCD (T5) inserts with a coolant
pressure of 7 MPa at a cutting speed of (a) 175 m min
-1
and (b) 200 m min
-1
.
195
xxi
Figure 4.75 - Surfaces generated after machining with PCD (T5) inserts with a coolant
pressure of 11 MPa at a cutting speed of (a) 200 m min
-1
and (b) 250 m min
-1
.
195
Figure 4.76 - Surfaces generated after machining with PCD (T5) inserts with a coolant
pressure of 20.3 MPa at a cutting speed of (a) 175 m min
-1
and (b) 200 m min
-1
.
196
Figure 4.77 - Surfaces generated after machining with PCD (T6) inserts with
conventional coolant supply at a cutting speed of (a) 175 m min
-1
and (b) 200 m min
-1
.
196
Figure 4.78 - Surfaces generated after machining with PCD (T6) inserts with a coolant
pressure of 11 MPa at a cutting speed of (a) 175 m min-1 and (b) 230 m min-1.
196
Figure 4.79 - Surfaces generated after machining with PCD (T6) inserts with a coolant
pressure of 20.3 MPa at a cutting speed of (a) 175 m min
-1
and (b) 200 m min
-1
.
197
Figure 4.80 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T5) insert
with conventional coolant supply.
197
Figure 4.81 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T5) insert
with 7 MPa coolant pressure supply.
198
Figure 4.82 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T5) insert
with 11 MPa coolant pressure supply.
198
Figure 4.83 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T5) insert
with 20.3 MPa coolant pressure supply.
199
Figure 4.84 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T6) insert
with conventional coolant supply.
199
Figure 4.85 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T6) insert
with 11 MPa coolant pressure supply.
200
Figure 4.86 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T6) insert
with 20.3 MPa coolant pressure supply.
200
Figure 4.87 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T5) insert
with conventional coolant supply at a cutting speed of (a) 140 m min
-1
and
(b) 230 m min
-1
.
201
Figure 4.88 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T5) insert
with a coolant pressure of 7 MPa at a cutting speed of (a) 175 m min
-1
and
(b) 250 m min
-1
.
201
Figure 4.89 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T5) insert
with a coolant pressure of 11 MPa at a cutting speed of (a) 175 m min
-1
and
(b) 230 m min
-1
.
202
Figure 4.90 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T5) insert
with a coolant pressure of 20.3 MPa at a cutting speed of (a) 200 m min
-1
and
(b) 230 m min
-1
.
202
Figure 4.91 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T6) insert
with conventional coolant supply at a cutting speed of (a) 140 m min
-1
and
(b) 200 m min
-1
.
202
Figure 4.92 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T6) insert
with a coolant pressure of 11 MPa at a cutting speed of (a) 175 m min
-1
and
(b) 230 m min
-1
.
203
Figure 4.93 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T6) insert
with a coolant pressure of 20.3 MPa at a cutting speed of (a) 175 m min
-1
and
(b) 230 m min
-1
.
203
Figure 4.94 - Chips generated when machining Ti-6Al-4V alloy with different grades of
PCD under various cutting conditions: (a): snarled chip; (e): long continuous chip, (b),
(c), (d), (f) and (g): segmented C-shaped chips.
204
Figure 4.95 - Tool life recorded when machining Ti-6Al-4V alloy with different CBN 205
xxii
tools (T6, T7 and T8) grades with conventional coolant flow (CCF), high coolant
pressures of 11 MPa and 20.3 MPa at various cutting speed conditions.
Figure 4.96 - Wear rate curves of different CBN tools when machining Ti-6Al-4V alloy
with conventional coolant flow and high pressures coolant supplies at various speed
conditions.
207
Figure 4.97 - Worn CBN 10 (T7) inserts after machining Ti-6Al-4V alloy using
conventional coolant supply at a speed of (a) 150 m min
-1
and (b) 200 m min
-1
.
207
Figure 4.98 - Worn CBN 10 (T7) inserts after machining Ti-6Al-4V alloy with 11 MPa
coolant pressure at a speed of (a) 150 m min
-1
and (b) 250 m min
-1
.
208
Figure 4.99 - Worn CBN 10 (T7) inserts after machining Ti-6Al-4V alloy with 20.3
MPa coolant pressure at a speed of (a) 150 m min
-1
and (b) 250 m min
-1
.
208
Figure 4.100 - Worn CBN 300 (T8) inserts after machining Ti-6Al-4V alloy using
conventional coolant supply at a speed of 150 m min
-1
.
209
Figure 4.101 - (a) Worn CBN 300 (T8) insert after machining Ti-6Al-4V alloy with 11
MPa coolant pressure at a speed of 150 m min
-1
and (b) enlarged section of worn
surface.
209
Figure 4.102 - Worn CBN 300 (T8) inserts machining Ti-6Al-4V alloy with 20.3 MPa
coolant pressure at a speed of (a) 200 m min
-1
and (b) 250 m min
-1
.
210
Figure 4.103 - Worn CBN 300-P (T9) inserts after machining Ti-6Al-4V alloy with
conventional coolant supply at a speed of (a) 150 m min
-1
and (b) with 11 MPa coolant
pressure at a speed of 250 m min
-1
.
210
Figure 4.104 - Worn CBN 300-P (T9) inserts after machining Ti-6Al-4V alloy with
20.3 MPa coolant pressure at a speed of (a) 150 m min
-1
and (b) 200 m min
-1
.
210
Figure 4.105 - Cutting forces (Fc) recorded at the beginning of cut when machining Ti-
6Al-4V alloy with different CBN inserts at various cutting conditions.
211
Figure 4.106 - Feed forces (Ff) recorded at the beginning of cut when machining Ti-6Al-
4V alloy with different CBN inserts at various cutting conditions.
212
Figure 4.107 - Surface roughness values recorded at the beginning of cut when
machining Ti-6Al-4V alloy with CBN inserts at various cutting conditions.
213
Figure 4.108 - Roundness variation recorded at end of cut when machining Ti-6Al-4V
alloy with CBN inserts using conventional coolant flow and high coolant supply
pressures at a speed of 150 m min
-1
.
213
Figure 4.109 - Hardness variation after machining Ti-6Al-4V alloy with CBN 10 (T7)
tools with conventional coolant flow and high coolant supply pressures at a speed of
150 m min
-1
.
214
Figure 4.110 - Hardness variation after machining Ti-6Al-4V alloy with CBN 300 (T8)
tools with conventional coolant flow and high coolant supply pressures at a speed of
150 m min
-1
.
214
Figure 4.111 - Hardness variation after machining Ti-6Al-4V alloy with CBN 300-P
(T9) tools with conventional coolant flow and high coolant supply pressures at a speed
of 150 m min
-1
.
215
Figure 4.112 - Microstructure of Ti-6Al-4V alloy after machining with CBN 10 (T7)
tools with (a) conventional coolant flow and (b) high coolant pressure of 20.3 MPa at a
cutting speed of 150 m min
-1
.
215
Figure 4.113 - Microstructure of Ti-6Al-4V alloy after machining with CBN 300 (T8)
tools at (a) 11 MPa and (b) 20.3 MPa coolant pressure at a cutting speed of
150 m min-1.
216
Figure 4.114 - Microstructure of Ti-6Al-4V alloy after machining with CBN 300-P (T9)
tools at (a) 11 MPa and (b) 20.3 MPa coolant pressure at a cutting speed of
216
xxiii
150 m min
-1
.
Figure 4.115 - Chips generated when machining Ti-6Al-4V alloy with CBN tools with
various coolant supplies at a cutting speed of 150 m min
-1
.
217
Figure 4.116 - Nose wear curves of silicon carbide (SiCw) whisker reinforced alumina
ceramic - rhomboid-shaped (T10) and square-shaped (T11) - inserts after machining Ti-
6Al-4V at various cutting conditions.
218
Figure 4.117 - Tool life recorded when machining Ti-6Al-4V alloy with silicon carbide
(SiCw) whisker reinforced alumina ceramic - rhomboid-shaped (T10) and square-shaped
(T11) - inserts at various cutting conditions.
219
Figure 4.118 - Wear observed on rhomboid-shaped SiCw alumina ceramic (T10 grade)
insert after machining Ti-6Al-4V alloy with conventional coolant supply at a speed of
(a) 140 m min
-1
; coolant pressure of 11 MPa at speeds of (b) 140 m min
-1
and (c) 400 m
min
-1
, coolant pressure of 20.3 MPa at a speed of (d) 140 m min
-1
, and in an argon
enriched environment at speeds of (e) 200 m min
-1
and (f) 400 m min
-1
.
221
Figure 4.119 - Wear observed on square-shaped SiCw alumina ceramic (T11 grade)
insert after machining Ti-6Al-4V alloy with conventional coolant supply at speeds of
(a): 130 m min
-1
and (b): 200 m min
-1
.
222
Figure 4.120 - Cutting forces (Fc) recorded at the beginning of cut when machining Ti-
6Al-4V alloy with SiCw alumina ceramic (T10 and T11) inserts at various cutting
conditions.
223
Figure 4.121 - Feed forces (Ff) recorded at the beginning of cut when machining Ti-6Al-
4V alloy with SiCw alumina ceramic inserts (T10 and T11) at various cutting
conditions.
223
Figure 4.122 - Surface roughness values recorded at the beginning of cut when
machining Ti-6Al-4V alloy with SiCw alumina ceramic inserts (T10 and T11) at various
cutting conditions.
224
Figure 4.123 - Hardness variation after machining Ti-6Al-4V alloy with rhomboid-
shaped SiCw alumina ceramic insert (T10 grade) at various environments and at a speed
of 140 m min
-1
.
225
Figure 4.124 - Microstructure of Ti-6Al-4V alloy after machining with SiCw alumina
ceramic tool (T10 grade) with: (a) conventional coolant flow, (b) high coolant pressure
of 11 MPa and (c) high coolant pressure of 20.3 MPa at a cutting speed of 140 m min
-1
.
226
Figure 4.125 - Chips generated when machining Ti-6Al-4V alloy with SiCw alumina
ceramic inserts: T10 grade at a cutting speed of 140 m min
-1
with: (a) conventional
coolant flow, (b) argon enriched environment, (c) coolant pressure of 11 MPa; T11
grade with conventional coolant flow at cutting speeds of: (d) 130 m min
-1
and
(e) 200 m min
-1
.
227
Figure 4.126 - Notch wear rate when machining Ti-6Al-4V alloy with T12 and T13
nano-grain size ceramics tools, Al
2
O
3
and Si
3
N
4
base respectively, with conventional
coolant flow and at various speed conditions
228
Figure 4.127 - Recorded tool life when machining Ti-6Al-4V alloy with nano-grain size
ceramics tools (T12 and T13) with conventional coolant flow and at various cutting
speeds.
228
Figure 4.128 - Wear observed on nano-ceramic tools after machining with Ti-6Al-4V
alloy at different cutting speeds: T12 (a: 130 m min
-1
), (b: 200 m min
-1
);
T13 (c: 110 m min
-1
) and (d: 200 m min
-1
).
229
Figure 4.129 - Component forces (cutting forces: Fc and feed forces: Ff) recorded at the
beginning of cut when machining Ti-6Al-4V alloy with T12 and T13 tools with
conventional coolant flow.
230
xxiv
Figure 4.130 - Surface roughness values at the beginning of cut when machining Ti-6Al-
4V alloy with nano-ceramic (T12 and T13) tools with conventional coolant flow.
231
Figure 4.131 - Roundness values recorded at the end of cut when machining Ti-6Al-4V
alloy cut with nano-ceramic (T12 and T13) tools with conventional coolant flow.
231
Figure 4.132 - Chips generated when machining Ti-6Al-4V alloy with nano-ceramic
tools: T12 (a): 110 m min
-1
, b: 130 m min
-1
and c: 200 m min
-1
); T13 (d: 200 m min
-1
)
232
Figure 5.1 - Variation of reactive component forces with coolant pressure supply before
machining Ti-6Al-4V alloy.
255
xxv
TABLES
Table 2.1 - Classification of aerospace materials (FIELD, 1968). 13
Table 2.2 - Application of aerospace alloys (FIELD, 1968). 14
Table 2.3 - Nominal chemical composition (wt.%) of various commercially available
pure-titanium and titanium-base alloys (ASM HANDBOOK, 1998).
18
Table 2.4 - Properties of Ti-6Al-4V alloy compared with a medium carbon steel, AISI
1045 (MACHADO; WALLBANK, 1990)
22
Table 2.5 - Softening points of tool materials (KRAMER, 1987). 56
Table 2.6 - Tool materials properties and cost (ABRÃO, 1995). 91
Table 3.1 - Nominal chemical composition of Ti-6Al-4V alloy (wt. %). 127
Table 3.2 - Physical properties of Ti-6Al-4V alloy. 127
Table 3.3 - Specification, chemical and mechanical properties of the cutting tool
materials used in the machining trials.
133
Table 3.4 - Mechanical properties and chemical composition (wt. %) of nano-ceramic
tools material (square shape inserts).
133
Table 3.5 - Summary of the experimental tests carried out when finish turning of
Ti-6Al-4V alloy at a constant feed rate of 0.15 mm rev-1 and a depth of cut of 0.5 mm.
136
Table 4.1 - Percentage improvement in tool life relative to conventional coolant supply
after machining Ti-6Al-4V alloy with different grades of carbides.
149
Table 4.2 - Percentage improvement in tool life relative to conventional coolant supply
after machining Ti-6Al-4V alloy with PCD inserts (STD and MM grades).
183
xxvi
LIST OF SYMBOLS
φ
Shear plane angle
ρ
Angle for chip friction
µ
Coefficient of friction
δ
Shear strain at the primary shear plane
σ
f
Normal stress component
τ
f
Shear stress component
σ
fma
Maximum normal stress
γ
n
Tool normal rake angle
α
n
Tool normal clearance angle
β
n
Tool normal wedge angle
ε
r
Tool included angle
λ
s
Tool cutting edge inclination angle
τ
st
Shear strength of chip material
Al
2
O
3
Alumina oxide
ANOVA
Analysis of Variance: a statistical assessment of sample data to decide if
differences exist between various groups of data.
C.I. Confidence Interval of 99% for distribution.
C
2
H
5
OH Ethanol vapour
CBN Cubic Nitride Boron
CCF Conventional Cooling Flow delivery system
CCl
4
Tetrachloromethane
DOC Depth of Cut (mm)
DOC Depth of Cut (mm)
EP Extreme Pressure
f Feed rate (mm rev
-1
)
HfC Hafnium carbide
HfN Hafnium nitride
HIPed Hot Isostatic Pressing
HPC High Pressure Coolant delivery system
HV Hardness Vickers
ISO International Standardisation Organisation
k
Constant for the work material
k´
r
Tool minor cutting edge angle
k
r
Tool cutting edge angle
m
meter
Max. Maximum value of a measured hardness
MgO Magnesium oxide
Min. Minimum value of a measured hardness
MM Multi-Modal Grade of Polycrystalline Diamond insert
MMC Metal Removal Rate (cm
3
min
-1
)
MQL Minimum Quantity Lubrication
n Spindle Speed (rev min
-1
)
PCD Polycrystalline Diamond
Pn Cutting edge normal plane
Pr Tool reference plane
xxvii
Pre Working reference plane
Ps Tool cutting edge plane
Pse Working cutting edge plane
r
Chip thickness ratio
Ra Average Surface Roughness
SIALON Silicon aluminium oxynitride
SiB
3
BNB
4
B Silicon nitride
STD Standard Grade of Polycrystalline Diamond insert
T1 Uncoated carbide insert – 883 designation
t
1
Underformed chip thickness
t
1
Actual chip thickness
T1 Uncoated carbide insert – 883 designation
T2 Uncoated carbide insert – 890 designation
T3 Coated carbide insert – CP 200 designation
T4 Coated carbide insert – CP 250 designation
T5 PCD insert – STD designation
T6 PCD insert – MM designation
T7 CBN insert – 10 designation
T8 Solid CBN insert –300 designation
T9 Solid Coated CBN insert – 300-P designation
T10
Silicon carbide whisker reinforced alumina ceramic insert – micron-grain size -
WG 300 designation – rhomboid-shaped geometry
T11
Silicon carbide whisker reinforced alumina ceramic insert – micron-grain size -
designation WG 300 – squared-shaped geometry
T12 Alumina base nano-grain size ceramic insert – SAZT2 designation
T13 Silicon nitride base nano-grain size ceramic insert – SNCTN1 designation
TaC Tantalum carbide
TD Theoretical Density
TiB
2
Titanium diboride
TiC Titanium carbide
TiN Titanium nitride
TiAlN Titanium aluminium nitride
TiZrN Titanium zirconium nitride
TiO
2
Titanium oxide
T
M
Cutting temperature
V
Cutting Speed (m min
-1
)
Vc
Chip Velocity (m min
-1
)
WC Tungsten carbide
Y
2
O
3
Yttria
ZrO
2
Zirconium oxide
CHAPTER I
INTRODUCTION
The machinability of titanium alloys is generally considered to be poor due to their
inherent properties such as chemical reactivity, consequently their tendency to weld onto the
cutting tool during machining leading to excessive chipping and/or premature tool failure. The
low thermal conductivity of titanium alloys increases temperature generated at the tool-
workpiece interface, adversely affecting tool life. They also exhibit tendency to form localised
shears bands (ASPINWALL et al., 2003) and work-harden during machining. Additionally,
their high strength maintained at elevated temperature and their low modulus of elasticity
further impair their machinability. These pose considerable problems in manufacturing hence
titanium-alloys have poor machinability (MILLER (1996), EZUGWU; WANG (1997),
VIGNEAU (1997), GATTO; IULIANO (1997)). The poor machinability of titanium alloys
have prompted many large companies (e.g. Rolls-Royce and General Electrics) to invest large
sums of money in developing techniques to minimise machining and overall processing costs
(EZUGWU; WANG, 1997). The best tool material is one that will maximise the efficiency
and ensure accuracy at the lowest cost, in other words, one that will satisfy the requirements
of a specific workpiece material (OKEKE, 1999). A cutting tool must possess high resistance
to abrasion in order to withstand changes in dimensions by rubbing action; hot-hardness to
maintain a sharp and consistent cutting edge when machining at elevated temperature
conditions; chemical stability (lack of affinity between the tool and workpiece) in order to
avoid the formation of a built-up edge; high resistance to thermal shock in order to withstand
continuous heating and cooling cycles (typical in milling operation) and high toughness which
allows the insert to absorb the forces and shock loads during machining. If a machine tool is
not sufficiently tough, then induced shock load alone can cause the edge to chatter.
2
Despite the developments in cutting tool materials for the machining of difficult-to-
machine materials at higher metal removal rates, they tend to be ineffective in machining
titanium-alloys because of their high chemical affinity. Also, recent developments in coating
technology seem to demonstrate only marginal improvement when machining titanium-alloys,
despite additional cost of the coated inserts. Ceramics and Cubic Boron Nitride
(CBN)/Polycrystalline Cubic Boron Nitride (PCBN) tools are not usually recommended for
machining titanium-alloys because of their poor performance due to excessive wear rates as a
result of the high reactivity of titanium-alloys to the tool materials in addition to their
relatively high cost (HONG; MARKUS; JEONG, 2001). Cutting tools used for machining
titanium alloys generally exhibit accelerated wear as a result of extreme thermal and
mechanical stresses close to the cutting edge. An ideal cutting tool for machining titanium
should have, among others, a hot hardness property to withstand elevated temperatures
generated at relatively high speed conditions. Reduction of hot hardness at elevated
temperature conditions lead to the weakening of the inter-particle bond strength and the
consequent acceleration of tool wear. In addition to that, the machining environment plays a
very important role in order to improve the machinability of titanium alloys.
Aero-engine alloys, particularly titanium alloys, cannot be effectively machined without
cooling. There is excessive concentration of temperature at the cutting interfaces when
machining titanium alloys because of their poor thermal conductivity. In addition to that,
practically all the energy consumed in machining is converted into thermal energy. Cutting
fluids are used to minimise problems associated with the high temperature and high stresses
generated at the cutting edge of the tool during machining. Titanium alloys are generally
machined using conventional coolant flow. Also, there is other technique to deliver coolant in
variable quantities at high/ultra high pressures, generally within the range 0.5 – 360 MPa
(SECO TOOLS (2002a)). This technique has been employed when machining mainly nickel
alloys. One of the benefits of using high pressure coolant supply is because it acts as a chip-
breaker. Additionally, the temperature gradient is reduced by penetration of the high-energy
jet into the tool-chip interface and consequently eliminating the seizure effect
(MAZURKIEWICZ; KUBALA; CHOW, 1989), thereby providing adequate lubrication at the
tool-chip interface with a significant reduction in friction (EZUGWU; BONNEY; YAMANE,
2003). These combined with high velocity coolant flow causes the breakage of the
continuous-type chips into very small segments. Because the tool-chip contact time is shorter,
the tool is less susceptible to dissolution wear caused by chemical reaction with newly
3
generated chips, especially titanium-alloy chips (LINDEKE; SCHOENIG; KHAN, 1991).
Increase in productivity has been noticed using high pressure coolant delivery relative to the
conventional methods of coolant delivery when machining nickel and titanium alloys at lower
speed conditions. Other cooling technique like the minimum quantity of lubrication (MQL)
has shown considerably improvement in the machinability of aerospace alloys compared to
conventional coolant flow and looks promising for machining titanium alloys in order to
improve the tribological processes present at the tool-workpiece interface and at the same
time eliminate environmental damages as well as minimizing some serious problems
regarding the health and safety of operators (SOKOVIC; MIJANOVIC (2001), DA SILVA;
BIANCHI (2000), LI et al. (2000), MACHADO; WALLBANK (1997)). With the same
purpose other environments such as atmospheric air (dry machining), argon enriched
environment and liquid nitrogen (cryogenic machining) are also been employed as alternative
cooling technology to improve the machinability of titanium-alloys. Since the gases can alter
the tribological conditions existing between two surfaces in contact such as the cutting zone
during machining, other environments such as atmospheres, dried air, oxygen, nitrogen, CO
2
and organic compounds such as tetrachloromethane (CCl
4
) and ethanol vapour (C
2
H
5
OH) are
also expected to improve the machinability of titanium-alloys. Some special machining
techniques including specially designed ledge tools, self-propelled rotary tool (SPRT),
ramping technique (taper turning) and hot machining have shown remarkable success in when
machining titanium alloys (EZUGWU; BONNEY; YAMANE (2003), EZUGWU; WANG
(1997), EZUGWU (2005)).
This thesis on the machining Ti-6Al-4V alloy with various cutting tools and different
cooling environments was developed in collaborative program with industrial partners: Rolls-
Royce Plc (aero-engine manufacturer), SECO Tools (cutting tool manufacturer) and Pumps
and Equipment Ltd (Warwick) who provided the high-pressure coolant delivery system for
this study. A comprehensive literature survey on the machinability of aero-engine alloys
under various cutting environments as well as the experimental techniques adopted in all
stages of the research programme such as turning tests, data acquisition, sample preparation,
analysis of the worn tools and machined surfaces, as well as initial machining results are
presented in this thesis. An investigation of the machinability of components manufactured
with titanium-base, Ti-6Al-4V (or IMI 318), alloy will involve the following:
4
i) Evaluation of recently developed cutting tools materials (uncoated and coated
cemented carbides, Polycrystalline Diamond (PCD) inserts, Cubic Boron Nitride (CBN) and
SiC Whiskers Reinforced Al
2
O
3
Ceramics) when machining titanium-base, Ti-6Al-4V, alloy
at high speed conditions;
ii) Cutting environments (high pressure coolant supplies at pressures of 7 MPa
(70 bar), 11 MPa (110 bar) and 20.3 MPa (203 bar), argon enriched environment, and
conventional coolant flow;
iii) Validation of the optimum machining conditions achieved on prototype component
without compromising its integrity.
1.1 Aims of the thesis
This thesis is geared primarily to achieve a step increase in the machining productivity
of a commercially available titanium-base, Ti-6Al-4V, alloy using recently developed cutting
tool materials, machining techniques and various cooling media such as conventional coolant
flow, high pressure coolant supplies and argon enriched environment. This study is part of the
Joint Strike Fighter (JSF) project – a vectored thrust, multi-role combat aircraft designed for
conventional take-off and landing or a Navy version which requires Short Take Off/Vertical
Landing capability in collaboration with Rolls-Royce plc. The thesis aims primarily towards
significant reduction in cost of manufacturing jet engines in the immediate future using
modern cutting tool technology and machining techniques.
The literature survey section covers cutting tool materials and the various cutting
environments employed in the machining of aero-engine alloys. The objectives of this thesis
are listed below:
Investigation of the effect of various cooling media (high-pressure coolant supply,
argon enriched environment and conventional coolant flow) on tool performance
when finish turning of titanium-base, Ti-6Al-4V (IMI 318), alloy;
Investigation of the dominant tool failure modes and wear mechanisms of newly
developed cutting tools (uncoated and coated cemented carbides, different grades of
Polycrystalline Diamond (PCD), Cubic Boron Nitride (CBN), SiC Whiskers
Reinforced Ceramic, and Al
2
O
3
and Si
3
N
4
base nano-grain ceramic inserts) when
finish turning of titanium-base, Ti-6Al-4V (IMI 318), alloy at high speed
machining;
5
Analysis of the surface finish and surface integrity of machined surfaces as well as
run-out of the machined bars;
Selection of the best combination of cutting tool-cutting environment-cutting
conditions to employ in the machining of prototypes/scaled down models of the 3
bearing swivel nozzle.
CHAPTER II
LITERATURE SURVEY
2.1 Historical Background of Machining
At the end of second ice age, more than 370,000 years ago, the mortal remains of a
recognizable human were preserved in a cave near Pekin. The presence of flakes of stone and
burned bones surrounding him was evidence that fire and tools had been discovered
(ARMYTAGE, 1970). Fire enabled him to live in cold countries, prepare food, frighten
animals away and lighten the gloom of his caves as well as use it to make pottery and stabilize
his life. Tools were no less important. As homo faber, man the maker, his first tools were
made of stone. By 3000 B.C. the communities cultivating the alluvial plains of the Nile, the
Tigris-Euphrates and the Indus valleys were producing foodstuff that they could use labourers
to dig canals, employ artisans to manufacture tools and support merchants who bought other
produce. They used sails to propel boats, oxen to draw ploughs, wheels to bear carriages and
metals to fabricate tools and facilitate arduous manual operations. The beginning of
civilization among people in Egypt (early third millennium B.C.) coincided with mining.
These mines were worked by slaves, obtained by war. Workers often used hammers and
wedges, bones and horns. At the beginning of the 2
nd
century before Christ, a greek called
Ctesibius compressed air to work a gun, built a water clock and a water organ sprayed fire
through a hose. Hero of Alexandria, in the 1
st
century B.C., constructed pneumatic devices
like water organs and ingenious toys of all kinds, including a steam turbine called an aeolipile.
He gave an account of five simple machines which become the basis of early technology: the
wedge, the pulley, the lever, the wheel and axle, and the endless screw. Metals obtained at
that time includes gold, silver, iron, lead, and later copper and tin were used in various ways.
7
Ethiopian kings bound their prisoners with gold chains. By about 1350 B.C. it is known that
Argonatus raided Colchis in Georgia looking for gold. Silver was used for handles on shields
(ARMYTAGE, 1970).
Machine tools, screw-cutting lathe and pumping plants were first described by Galileo´s
disciple in 1579, professor Jacques Besson (Orleans), in his Theâtre des Instruments
Mathematiques et Méchaniques. However, it was only from 18
th
century that great technical
advances in machine tools began. The early developments were laying the base for industrial
development. At the beginning of the 18
th
century, wood was the dominant workpiece
material and the machining of metal was very limited and quite crude. By about 1740, the
power revolution in Britain, and the development of the steam engine, led to the development
of superior engineering materials like, cast iron, wrought iron, bronze and brass which were
relatively easy to machine with tools materials available in that time: the carbon tools steels
hardened and tempered by blacksmiths (TRENT; WRIGHT, 2000). There were early planning
and milling machines as well as lathes that could perform threading. The introduction of the
cross-slide on a lathe represented a great progress, meaning that tools did not need to be held
by hand, they could be secured in a tool-post instead. Machining operation was very slow at
the beginning of the 19
th
century, one example being the shapping of one and half square
metres of an iron surface taking all of one working day. The cutting depth and the length of
stroke were set and the shapping machine was left to run. The development of workshops and
their machinery was extensive in Europe and in America during 19
th
century. They were
mainly stimulated by the armament industry, navigation and railway industries. On the
American continent, arms makers led the progress and developed machines and introduced
manufacturing based on interchangeable parts and standardised measurements. The turret
lathe was introduced as a major quick-change of tools and turret lathes and automatic screw
machines were widely in operation by the end of the 19
th
century in industrialised countries
(AB SANDVIK COROMANT, 1994).
From the 1860s, the expanding economy promoted increase in productivity. With
further development of new metals and alloys (considered more difficult to machine) like
steel for instance, even the best carbon steel tools were pushed to their functional limits, thus
becoming insufficient for manufacturers’ needs, constraining production speed and hampering
efficiency. It was at this stage that emphasis was shifted from the development of basic
machine tools to cutting tool materials which could withstand the severe conditions in metal
cutting i.e. tool materials which could cut at higher cutting conditions. However, by the end of
8
the 19
th
century both labour and capital costs of machining increased significantly. Reducing
costs by accelerating and automating the cutting process became necessary. Machine tools
with manual control have became uneconomical as the required output rises, and the need to
design and produce advanced machines became indispensable. At the same time, advances in
technology led to developments in metal cutting field. Production of machine tools in order to
maximise the utilisation of each generation of the cutting tool materials has been created.
Semi-automatic and automatic machines became available. Additionally, the engineering and
economic requirements of the performance of machined components were constantly
changing and becoming more demanding. Technical problems necessitated a more realistic
appraisal of mechanical engineering design problems. The designer must critically examine
his selection of material design products, strength properties, processing operations, and
means for securing adequate material inspection (ASME HANDBOOK, 1965).
In the 20
th
century the introduction of automatic machines, computer numerically
controlled (CNC) machines and transfer machines allowed better tool efficiency to be
achieved. Subsequently the development of the jet engine by Great Britain, United States and
German, forced mechanical and metallurgical engineers to work towards the improvement
and/or development of tool materials capable of withstanding the demands imposed by the
new materials. Optimisation of shape and geometry of tools, as well as changing the
properties of cutting tools by manipulation of chemical composition have been made to
prolong tool life at higher cutting speeds; development of new coolants and lubricants ensured
to improve surface finish and higher rates of metal removal. Advances in materials science
enabled the development of novel materials that can operate under aggressive environments
such as severe conditions of stress, high temperature and corrosion. Aluminium-base alloys,
magnesium-base alloys, high-alloy steels, nickel-base and titanium-base alloys and other
metal matrix composites are examples of these materials currently in use. They led to the
evolution of new cutting tool materials capable of providing long tool life, high stock removal
rates and greater component accuracy and integrity.
2.2 Overview of Aerospace Alloys
Living in the modern society will involve the use of manufactured items such as
telephones for communication, cars and aeroplanes for transportation, etc. It therefore
becomes unavoidable to stop the changes or trends taking place in the modern world. There is
9
a need to adjust to this situation, particularly in cases where demand for better quality and
reliability are critical factors that have led companies to design the process technology and to
ensure that the system is flexible, controllable, and as efficient as possible in order to remain
competitive in the global market place (DE GARMO; BLACK; KOHSER, 1999). In this
context, rapid advances in aircraft and space vehicles have created a need for the development
of many new materials to withstand extreme temperatures, high stresses and unusual
environments. Aircraft and space vehicle structures operate at elevated temperatures and
stresses than formerly served by aluminium and magnesium alloys. High strength steels, high
temperature alloys, refractory alloys, beryllium, and plastics are materials in common use in
the aerospace industry. Most of the aerospace alloys are susceptible to surface alterations or
damage during machining, hence extreme care must be exercised to maintain surface integrity
of machined components (FIELD, 1968). Apart from continuous efforts to improve engine
performance, weight reduction and manufacturing costs also come into play, indeed
increasingly so. The best preparation for the intense competitive pricing in the aerospace
industry is by reducing overall manufacturing cost (MILLER, 1996).
2.2.1 Aero-Engine Alloys
The aero-engines, also called gas turbine engines or jet engines, are responsible for
propelling most modern passenger and military aircrafts. An aero-engine consists of an
inlet,
compressor, shaft, burner and nozzle as are shown in Figures 2.1 (a) and (b) (BENSON
(2002), ROLLS-ROYCE PLC (2003)). High performance, light weigh, low life cycle cost are
the main requirements for aero-engines in addition to the need for increasing their thrust-to-
weight ratio (T/W). A way to increase this performance is by increasing turbine inlet
temperature (TIT). The TIT increased significantly from the 1980’s onwards. Figure 2.2
shows the trend in turbine inlet temperature in aero-engines (OHNABE et al., 1999). Figure
2.3 shows that some improvements in the performance of aero-engine alloys can be associated
with ability of the materials designer to provide the engineer with materials that can withstand
high temperatures conditions (BENSON, 2002). Figure 2.4 shows the maximum service
temperature chart to help with identification of new possibilities for materials development
(LOVATT; SHERCLIFF, 2002).
10
(a)
(b)
Figure 2.1 - (a) A typical jet engine and its main parts (Pratt and Whitney F100 jet engine)
(BENSON, 2002); (b) typical jet engine (Trent 700) manufactured by Rolls-Royce plc
(ROLLS-ROYCE PLC, 2003).
11
Figure 2.2 - Trends in turbine inlet temperature in areo-engines (OHNABE et al., 1999).
Figure 2.3 - Improvements in aero-engine performance (BENSON, 2002).
Carbon composites have been used in the aero-engine industry since the 1970s in
relatively smaller quantities. They are currently attracting more attention due to their
improved properties and manufacturing processes. Figure 2.5 shows trend of materials usage
in typical aero-engines. Increasing demand for titanium and nickel-base alloys up till the end
of the 20
th
century is clearly illustrated, highlighting their dominant and competitive use in the
manufacture of aerospace engines (MILLER, 1996).
Materials employed in the aerospace industry can be grouped into different categories
according to their nominal composition, Table 2.1 (FIELD, 1968). General areas of
application of various categories of materials used on the aerospace industry are described in
Table 2.2 (FIELD, 1968).
12
Figure 2.4 - Maximum service temperature of various materials (LOVATT; SHERCLIFF,
2002).
.
Figure 2.5 - Trend of materials usage in aero-engines (MILLER, 1996).
13
Table 2.1 - Classification of aerospace materials (FIELD, 1968).
Type of material Designation Nominal composition
Aluminium alloy 2024 Al-4.5Cu-1.5Mg
Austenitic stainless steels 302 Fe-18Cr-9Ni
321 Fe-18Cr-11Ni
Martensitic stainless steel 410 Fe-12Cr
Precipitation-hardening stainless
steels
17-7 PH Fe-17Cr-7Ni-1Al
17-4 PH Fe-16.5Cr-4Ni-4Cu
High-strength steels 4340 Fe-0.8Cr-1.8Ni-0.4C
H-11 Fe-5.25Cr-1.35Mo-0.5V-0.37C
D6AC Fe-1Cr-0.5Ni-1Mo-0.46C
18% Ni maraging
steel
Fe-18Cr-8.5Co-3.25Mo
Nickel-base high-temperature alloys Inconel 718 Ni-19Cr-3Mo-5.2Cb-0.8Ti-0.6Al-18Fe
Udimet 500 Ni-18Cr-16Co-4Mo-2.8Ti-2.8Al
Rene 41 Ni-19Cr-11Co-10Mo-3.1Ti-1.5Al
Waspaloy Ni-19.5Cr-13.5Co-4.2Mo-3Ti-1.2Al-2Fe
Cobalt-base high temperature alloys HS-25 Co-20Cr-10Ni-15W-3Fe
Iron-base high temperature alloy A-286 Fe-15Cr-26Ni-1.2Mo-2Ti
Titanium alloys Ti-6Al-4V
(IMI 318)
Ti-6Al-4V
Ti-8Al-1Mo-1V Ti-8Al-1Mo-1V
Ti-3Al-13V-11Cr Ti-3Al-13V-11Cr
Pure tungsten Unalloyed tungsten W plus trace elements
Molybdenum alloy TZM Mo-0.5Ti-0.07Zr
Columbium alloy Columbium D31 Cb-10Mo-10Ti
Tantalum alloy Ta-10W Ta-10W
Beryllium I-400 Be-4.25BeO
Composite and plastics Fibre reinforced
plastics
-----
14
Table 2.2 - Application of aerospace alloys (FIELD, 1968).
Material Application
Austenitic stainless steel Compressor rings and housings, wing
panels, cryogenic vessels
Martensitic stainless steel Compressor blades, disks and rings and
Precipitation-hardening stainless steels spars, struts, fittings, fasteners, wing panels,
High-strength steels Compressors disks, airplane spars, rocket
motor cases, longerons, struts, bulkheads,
landing gear, fitting
Nickel-or colbalt-base high-temperature
alloys
Turbine disks, turbine blades
Titanium alloys Airplane spars, wing skins and panels,
longerons, struts, bulkheads, fasteners,
rocket motor cases
Refractory alloys Space propulsion structures, nozzles and
nozzles inserts
Beryllium Compressors disks and blades, re-entry
capsules, guidance systems
Composites and plastics Nose cones, thermal barriers, heat shields,
ablative components, compressor blades,
rocket motor cases
2.3 Superalloys
The origin of the term “Superalloy” refers to “Heat Resisting Alloys” or “High
Temperature Alloys”. The nomenclature “superalloy” became popular in the late 1940s, after
Word War II, to describe a group of alloys developed for use in turbo superchargers and
aircraft turbine engines which operated at elevated temperature service in severe corrosive
environments where relatively severe mechanical stressing is encountered and where surface
stability frequently is required. These alloys, usually based on Group VIII A elements, consist
of various formulations made from the following elements: nickel, iron, cobalt and chromium,
as well as lesser amounts of niobium, tantalum, titanium, tungsten, molybdenum and
aluminium. The most important properties of the superalloys are long-time strength at
temperatures above 650 ºC (1200 ºF) and resistance to hot corrosion and erosion (ASM
INTERNATIONAL, 1988).
Many alloys are used at elevated temperatures and must be able to withstand the
deteriorating effect of the service atmosphere, as well as possess sufficient strength for the
design condition and have adequate stability to withstand damaging metallurgical structural
15
changes at operating temperature. For use at moderate temperatures (< 540 ºC) and moderate
stress, the 12% Cr corrosion-resistant steels are satisfactory. A group containing chromium
and molybdenum and/or carbide formers and/or cobalt or nickel, the so-called super 12% Cr
steels, have been used where conditions of higher stress and moderate temperature are
required. As the temperature is increased, under conditions of low stress, higher chromium
steels, either the ferritic corrosion-resistant or the austenitic stainless steels containing nickel
as well as chromium, or nickel-chromium alloys are employed. For very high operating
temperatures, there is increasing interest in the refractory metals of Groups V (vanadium,
niobium, and tantalum) and VI (chromium, molybdenum and tungsten) as well as ceramics.
The refractory metals, however, exhibit very poor oxidation resistance, and their use is
generally restricted to non-oxidizing environments. Because of low toughness for most
structural applications, ceramic materials have limited use. The high-temperature applications
of superalloys are extensive, including components for aircraft, chemical plant equipment and
petrochemical equipment.
2.3.1 Titanium Superalloys in the Aerospace Industry
Although all superalloys materials play a significant role in Aerospace industry, nickel-
base and titanium alloys are the major materials used due to their improved properties. In this
thesis the emphasis will be placed on the machinability of titanium-base, Ti-6Al-4V, alloy.
Wilhelm J. Kroll, in the late 1930s, was the person responsible for development of a
safe and economical method of production of titanium (ASM HANDBOOK, 1998). Initially,
Kroll´s process involved reduction of titanium tetrachloride (TiCl4) with sodium and calcium.
Later magnesium was utilized in an inert gas atmosphere. Research continued through World
War II and by the late 1940s, with definition of physical and mechanical properties as well as
alloying characteristics, that titanium really gained industrial and academic importance.
Titanium is encountered in two crystallographic forms. At room temperature, unalloyed
(commercially pure) titanium has a hexagonal close-packed (hcp) structure referred to as
alpha (α) phase. At 883
o
C (1621 ºF), titanium undergoes an allotropic transformation from
hcp to a body-centred cubic (bcc) structure known as beta (β) phase, which remains stable to
the melting point. The manipulation of these crystallographic variations through alloying
additions and thermomechanical processing is the basis for the development of a wide range
of alloys and properties. These phases also provide a convenient way to categorise titanium
mill products (ASM HANDBOOK, 1998). Based on these phases, titanium alloys can be
16
categorised into four main groups namely alpha (α), near alpha, alpha + beta and beta (β)
based on their atomic crystal structure. The alloying elements can be categorised according to
their effect on the stability of the α- and β-phases. Those that raise the transformation
temperature are known as α-stabilisers (Al, O, N and Ga) and those that decrease the
transformation temperature are called β-stabilisers (Mo, V, W and Ta). Cu, Mn, Fe, Ni, Co
and H are also β-stabilisers but form the eutectoid (BHADESHIA, 2003). Al is a very
effective α-strengthening element at ambient and elevated temperatures up to 550ºC and its
low density is an additional advantage. The residual elements (or impurities) such as C, N, Si
and Fe raise the strength and lower the ductility of titanium products. Basically, O and Fe
contents determine strength levels of commercially pure titanium. In higher strength grades, O
and Fe are intentionally added to the residual amounts already in the sponge to provide extra
strength. On the other hand, C and N usually are held to minimum residual levels to avoid
embrittlement. When good ductility and toughness are required, the extra-low interstitial
(ELI) grades are recommended (ASM HANDBOOK, 1998). Table 2.3 contains the chemical
composition by weight of commercially available pure-titanium and titanium-base alloys
(ASM HANDBOOK, 1998). Detailed description of titanium alloys and their properties have
been reported elsewhere (ASM HANDBOOK (1998), BOYER (1996), EZUGWU; WANG,
1997).
• Alpha alloy: represents the group of the commercially pure titanium alloy grade
with oxygen and iron as the primary alloying elements; Alpha alloys generally have
creep resistance superior to β alloys, and are preferred for high-temperature
applications. The absence of a ductile-to-brittle transition, a feature of β-alloys,
makes α-alloys suitable for cryogenic applications. Materials of this group are
characterised by satisfactory strength, toughness and weldabillity, but poorer
forgeability than β-alloys. A single phase α-alloy, Ti 5-2
2
1
(Ti-AL-2
2
1
Sn) is still
available commercially besides commercially-pure titanium;
• Near alpha alloy: contain large percentage of α-stabilisers and small amount of β -
stabilisers like molybdenum and vanadium. Typical alloys in this group include Ti-
3Al-2.5 V (Ti-3-2.5), Ti-8Al-1Mo-1V (Ti-8-1-1) and Ti-6Al-2Sn-4Zr-2Mo (Ti-6-2-
4-2S). These alloys have mainly used at operating temperatures between 400 to
520
o
C EZUGWU; WANG (1997);
17
• Alpha + beta alloy: this is a mixture of α- and β-phases and may contain between
10 and 50% β-phases at room temperature. The most common α + β alloys are Ti-
6Al-4V (Ti-6-4 or IMI 318), Ti-6Al-2Sn-4Zr-6Mo (Ti-6-2-4-6) and Ti-6Al-2Sn
alloys. The properties of these alloys can be controlled through heat treatment,
which is used to adjust the amounts and types of β-phase element. A good
combination of their properties ensure better operations within the temperature
range of 315-400
o
C (ASM HANDBOOK (1998). Solution treatment followed by
aging at 480 to 650 ºC (900 to 1200 ºF) precipitates α and β in a matrix of retained
or transformed β-phase;
• Beta alloy: these are referred to as high hardenability titanium alloys due to the β-
stabilisers, represented by Ti-15V-3Cr-3Al-3Sn (Ti-15-3), Timetal 21S (Ti-15Mo-
2.7Nb-3Al-0.2Si), Ti-3Al-8V-6Cr-4Mo-4Zr and Ti-10V-2Fe-3Al alloys. As stated
earlier, the transition elements present in these group such as Mo, V and Nb tend to
decrease the temperature of the α- to β-phase transition and thus promote
development of the bcc β-phase. They have excellent forgeability over a wider
range of forging temperatures than α-alloys (ASM HANDBOOK, 1998). They
also exhibit high stress corrosion resistance, can be heat-treated to high strengths
and offer fabrication advantages, particularly for producing sheets, owing to their
cold rolling capabilities. Some β-alloys, such as Ti-10-2-3 and β-C, have excellent
fatigue properties while others such as Ti-15-3 have, in general, poor fatigue
properties relative to their strengths.
18
Table 2.3 - Nominal chemical composition (wt.%) of various commercially available pure-
titanium and titanium-base alloys (BOYER, 1996).
Product Specification C H O N Fe Al Sn Zr Mo Others
JIS class 1
…
0.015 0.15 0.05 0.20 … … … … …
JIS class 2
…
0.015 0.20 0.05 0.25 … … … … …
JIS class 3
…
0.015 0.30 0.07 0.30 … … … … …
DIN 3.7025 0.08 0.013 0.10 0.05 0.20 … … … … …
DIN 3.7035 0.08 0.013 0.20 0.06 0.25 … … … … …
DIN 3.7055 0.10 0.013 0.25 0.06 0.35 … … … … …
GOST BT1-00 0.05 0.008 0.10 0.04 0.20 … … … … …
Pure titanium
ASTM grade 12
(UNS R53400)
0.10 0.015 0.25 0.03 0.30 … … … … …
Ti-2.5Cu 0.08 0.01 0.2 0.05 0.2 … … … …
2.0-
3.0Cu
Ti-5Al-2.5Sn
(DIN 17851)
0.08 0.02 0.2 0.05 0.5
4.0-
6.0
2.0-
3.0
… … ...
Ti-5Al-2.5Sn- ELI
(AMS 4909)
0.05 0.0125 0.2 0.035 0.25
4.5-
5.75
2.0-
3.0
… …
O +
Fé=0.32,
0.005Y
Ti-8Al-1V-1Mo
(MIL-R-81588)
0.035 0.005 0.12 0.015 0.20
7.35-
8.35
… …
0.75-
1.25
0.75-
1.25V
Ti-6242 0.05 0.0125 0.15 0.05 0.25
0.1Si,
0.005Y
Ti-679 0.04 0.008 0.17 0.04 0.12 2.25 11 5 1 0.2Si
Ti-6Al-2Sn-1.5Zr-
1Mo
... ... ... ... ... 6 2 1.5 1
0.35Bi,
0.1Si
IMI 829 ... ... ... ... ... 5.5 3.5 3 0.25
1Nb,
0.3Si
Ti-6Al-4V (IMI 318) 0.08 0.01 0.2 0.05
5.5-
6.75
5.5-
6.75
... ... ...
3.5-4.5V
0.005Y
Ti-6Al-4-ELI
(AMS 4907)
0.08 0.0125 0.13 0.05 0.25
5.5-
6.75
... ... ... 3.5-4.5V
Ti-6Al-6V-2Sn 0.05 0.015 0.20 0.04
0.35-
1.0
6 2 ... ...
0.75Cu,
6V
Ti6246 0.04 0.0125 0.15 0.04 0.15 6 2 4 6 ...
IMI 679 ... … … … … 2 11 4 1 0.25Si
Ti-13V-11Cr-3Al
(MIL-T-9047)
0.05 0.025 0.17 0.05 0.35
2.5-
3.5
... ... ...
12.5-
14.5V,
10-12Cr
Ti-8Mo-8V-2Fe-3Al 0.05 0.15 0.16 0.05
1.6-
2.4
2.6-
3.4
... ... 7.5-8.5 7.5-8.5V
Ti-10V-2Fe-3Al 0.05 0.015 0.13 0.05
1.6-
2.5
2.5-
3.5
... ... ...
9.25-
10.75V
Titanium base alloys
Transage 175 0.08 0.015 0.15 0.05 0.20
2.2-
3.2
6.5-
7.5
1.5-
2.5
… 12-14V
Titanium alloys possess high strengths, high strength-to-weight ratio relative to other
materials such as steel (60% density of steel) and high compatibility with composite structure
which can often bridge the properties gap between aluminium and steel alloys, providing
many of the desirable properties of each. For example, titanium like aluminium is non-
magnetic and has good heat-transfer properties (despite its relatively low thermal
19
conductivity). Thermal expansion coefficient of titanium alloys, ranging from about 9 to
11 ppm ºC
-1
, is slightly lower than that of most steels and less than half that of aluminium
(ASM HANDBOOK, 1998). Titanium could also replace aluminium when operating
temperature exceeds about 130ºC, the normal maximum operating temperature for
conventional aluminium (BOYER, 1996). Other important characteristics of titanium alloys
depend on the class of alloy and the morphology of the alpha constituents. In the near alpha
and alpha-beta alloys, the variations in the alpha morphology are achieved with different heat
treatments. Titanium alloys have high-temperature strength that is associated with the alpha
and near-alpha alloys. However, when creep strength is not a factor in an elevated-
temperature application, the short time elevated-temperature tensile strengths of beta alloys
have a distinct advantage.
Excellent corrosion resistance is another important property of titanium alloys.
Although titanium is a highly reactive metal, it also has an extremely high affinity for oxygen
and thus forms a very stable and highly adherent protective oxide film on its surface. This
oxide film forms spontaneously and instantly when fresh metal surfaces are exposed to air
and/or moisture. In addition to its passive behaviour, titanium is non-toxic and biologically
compatible, making it useful in applications ranging from chemical processing equipment to
surgical implants and prosthetic devices (ASM HANDBOOK, 1998). Its corrosion resistance
is such that corrosion protective coatings of paint are not required. Typical applications in
very corrosive environments which dictate the use of titanium to provide high structural
durability are the floor structure under the galleys and lavatories. Titanium alloys are also
predominantly used within medical and aerospace industries.
In biomedical applications, the effectiveness and reliability of titanium implants in
human body are essential requirements because, once they are installed, they cannot readily
be maintained or replaced. The natural selection of titanium for implantation is determined by
a combination of most favourable characteristics including immunity to corrosion,
biocompatibility, strength, low modulus and density and the capacity for joining with bone
and other tissue – osseointegration. The mechanical and physical properties of titanium alloys
combine to provide implants which are highly damage tolerant. A titanium implant has a
stiffness of less than half that of stainless steel or cobalt chrome, therefore reducing the effects
of weight shielding. Most metals in body fluids and tissue are found in stable organic
complexes. Titanium is judged to be completely inert and immune to corrosion by all body
fluids and tissue and is thus wholly biocompatible. Typical application of titanium in
20
biomedical area are: dental implants, bone and joint replacement, maxilla and cranium/facial
treatments, external prostheses, cardiovascular devices and surgical instruments (plates, pins
and hipjoints)
(TIG, 2002). Figure 2.6 (a) shows typical medical/dental applications of
titanium alloys
(TIG, 2002).
(a)
(b)
(c)
Figure 2.6 - Typical applications of titanium: (a) modular femoral components (prostheses
manufactured in titanium-base, Ti-6Al-4V, alloy, (b) valves and (c) screw (TIG, 2002).
Titanium industry market is mainly shared by commercial and military jet aircraft. Its
dependence on the aerospace industry, which is cyclical in nature, resulted in several setbacks
(Figure 2.7).
Figure 2.7 - Phases of titanium product life cycle in the U.S. (ASM HANDBOOK, 1998).
This Figure shows the rapid growth of the U.S. titanium industry as well as the phases
of titanium product life cycle. Despite continued development of new alloys and product
forms, titanium has moved rapidly through its product life cycle to maturity in the aircraft
21
industry. Titanium is still in growth stage in applications where corrosion resistant is required,
such as the marine and biomedical industries. Due to their low strength-to-weight ratio, tensile
strength up to 1200 MPa and capability of operating at temperatures from sub zero to 900°C
(ASM HANDBOOK, 1998), titanium alloys are widely used for highly loaded aerospace
components that operate at low to moderately elevated temperatures, including both airframe
(fittings, bolts, wing boxes, fuselage frames an panels, brake assemblies, hydraulic tubing etc)
and aero-engine components (discs, blades, shafts, afterburner cowlings, bolts, hot-air ducts
casings from the front fan to the last stage of the high pressure compressor, and at the rear end
of the engine for lightly loaded fabrications such as plug and nozzle assemblies). Polymer
matrix composite (PMC) compatibility has also becoming an important issue with higher
utilization of composite structure on aircraft. The titanium is galvanically compatible with the
carbon fibres in the composites, whereas aluminium (and low alloy steels) and carbon
generate a significant galvanic potential. The selection of titanium for this application is
related to the critically of the structure. There are corrosion protection systems which are used
to isolate aluminium from carbon composites to preclude the corrosion problem, but the
integrity of the coating over the life of the airframe must be taken into account (BOYER,
1996).
Other applications of titanium alloys are in the automotive industry (valve and
suspension assemblies, cam belt wheels, steering gear, connecting rods, drive shafts,
crankshafts, high strength fasteners) (Figure 2.6 (b) and (c)); construction/architecture (facing
and roofing, concrete reinforcement, monument refurbishment); chemical processing (storage
tanks, pumps, frames, pressurized reactors, piping and tubing); marine equipment (deep-sea
pressure hulls, submarines, submarine ball valves, data logging equipment, lifeboat parts,
cathodic protection anodes); machine tools (flexible tube connections, protective tubing,
instrumentation and control equipment); cutting implements (scissors, pliers, knives) and
consumer products (jewellery, watches, camera shutters, bicycle frames, tennis rackets, horse
shoes, target pistols).
The most widely used titanium alloy is the Ti-6Al-4V alpha-beta alloy. This alloy was
first introduced in 1954 and nowadays accounts for about 60% of the total titanium
production. It is processed to provide mill-annealed or β-annealed structures and is sometimes
solution treated and aged. Ti-6Al-4V is extensively used in the aerospace industry, in all mill
product forms, for airframe structural components (80%-90%) and other applications
requiring strength at temperatures up to 315ºC, because of its outstanding strength-to-weight
22
ratio relative to other materials (BOYER (1996), ASM HANDBOOK (1998)). Other
applications are for high-strength prosthetic implants and chemical-processing equipment. Its
tensile strength can reach up to 1100 MPa (ASM HANDBOOK, 1998). Ti-6Al-4V-ELI
(extra-low interstitial) alloy possesses reduced interstitial impurities, which improves ductility
and toughness. This grade is designed for cryogenic applications and fracture-critical
aerospace applications. Ti-6Al-4V alloy has limited section size hardenability hence is most
common used in the annealed condition. This alloy is very forgiving with variations in
fabrication operations, despite its relatively poor room-temperature shaping and forming
characteristics when compared to steel and aluminium. It also has useful creep resistance up
to 300ºC and excellent fatigue strength and fair weldability (ASM HANDBOOK, 1998). As a
basis for comparison, some of important properties of Ti-6Al-4V alloy and of AISI 1045 steel
are given in Table 2.4(MACHADO; WALLBANK, 1990).
Table 2.4 - Properties of Ti-6Al-4V alloy compared with a medium carbon steel, AISI 1045
(MACHADO; WALLBANK, 1990).
Material Tensile
Strength
(MPa)
Yield
Strength
(MPa)
Thermal
conductivity
(W m
-1
K
-1
)
Specific
heat at 20-
100ºC (J
kg
-1
K
-1
Hardness
(HV)
Density
(g cm
-3
)
Elongation
(%)
Reduction
in area (%)
Modulus of
elasticity
tension
(GPa)
Ti6Al4V
annealed
bar
895 825 7.3 580 340 4.43 10 20 110
Ti6Al4V
solution
treated
and aged
bar
1035 965 7.5 … 360 … 8 20 …
AISI
1045
cold
drawn
625 530 50.7 486 179 7.84 12 35 207
Other titanium alloys are designed for particular application areas. They are: Ti-5Al-
2Sn-2Zr-4Mo-4Cr (Ti-17) and Ti-6Al-2Sn-4Zr-6Mo designed for high strength in heavy
sections at elevated temperatures; Ti-6242S, IMI 829 and Ti-6242 designed for creep
resistance; Ti-6Al-2Nb-1Ta-1Mo and Ti-6Al-4V-ELI designed to resist stress corrosion in
aqueous salt solutions and for high fracture toughness; Ti-5Al-2.5Sn and the grade ELI
designed for weldability and cryogenic applications, respectively; Ti-6Al-6V-2Sn, Ti-10V-
23
2Fe-3Al designed for high strength at low-to-moderate temperatures (ASM HANDBOOK,
1998).
Due to the expensive cost of titanium alloys, relative to other metals, owing to the
complexity of the extraction process, difficulty of melting and problems during fabrication
and machining, integrated researches have been established in many countries to improve the
machinability of titanium alloys. Success in the machining of aerospace alloys, especially
titanium alloys, will depend on the correct selection of cutting tool (s), cutting environment
and appropriate cutting conditions for each machining operation (NORTH, 1986).
2.4 Machining Operations
Machining can be defined as a process in which parts with desired shape, size and finish
are obtained by removing unnecessary or excess material in form of swarf or chips (GHOSH;
MALLIK, 1986). Machining process is perhaps the most versatile manufacturing process.
Figure 2.8 illustrates schematically the main components involved in a conventional
machining operation (cylindrical machining or turning). It is common to use the term “metal
cutting” to refer to the machining process. Modern metal cutting produces many chips and
controlled chip formation is a prerequisite for any operation of metal removed (WATERS,
2000).
Figure 2.8 - Metal cutting diagram (WATERS, 2000).
24
Machine tools are generally designed for specific machining operations based on
features that are mainly related to how the relative motions are generated and on the cutting
tools employed. According to the British Standard Institution Publication (BRITISH
STANDARD, 1972), the important motions in machining are:
1. Primary Motion: also referred to as cutting motion. It is responsible for the cutting
action between the tool and the workpiece;
2. Secondary Motion: this motion is responsible for gradually feeding the uncut
portion, and thus also known as feed motion. This in addition to the primary motion,
leads to repeated or continuous chip removal and the creation of a machined surface
with the desired characteristics. This motion may proceed continuously or by steps;
3. Resultant Motion: motion resulting from simultaneous primary and secondary
motion.
2.4.1 Terminology used in Metal Cutting
The following useful terminologies are used to describe metal cutting operations
(BRITISH STANDARD, 1972) and illustrated in Figure 2.9 (KALPAKJIAN; SCHMID,
2000):
• Cutting Speed (V): is defined as the rate at which the surface of workpiece being
machined moves across the cutting edge of the tool and measured in metres per
minute (m min
-1
). The following formula, Equation (2.1), is used to calculate the
cutting speed:
1000
nd
V
××
=
π
(2.1)
Where d is the diameter of the workpiece material
n is the spindle speed expressed in revolutions per minute (rev min
-1
);
• Feed Rate (f): is defined as the distance that the tool travels in an axial direction
during each revolution of the workpiece material and expressed in millimetres per
revolution (mm rev
-1
);
• Depth of Cut or width of cut (DOC or w): is defined as the thickness of the
unwanted material removed from the workpiece in the radial direction and
expressed in millimetres (mm);
• Work Surface: the surface on the workpiece to be removed by machining operation;
25
• Machined Surface: the desired surface formed on the workpiece which is produced
by the action of the cutting tool;
• Transient Surface: the part of the surface of the workpiece which is formed by the
cutting edge and removed during the following cutting revolution or stroke of the
tool or workpiece material;
• Metal Removal Rate (MRR): this is the product of the cutting speed, feed rate and
the depth of cut, and is a parameter often used to determine the efficiency of a
machining process given by the Equation (2.2). It is expressed in cm3 min-1:
• MRR = V× f × DOC • (2.2)
• Tool Life Line (L): is the product of the cutting speed and tool life (T). This is the
parameter that can determine the optimum cutting speed when machining at a given
feed rate and depth of cut. The tool life line is calculated by the Equation (2.3) and
is expressed in metres (m):
L = VxT
(2.3)
Figure 2.9 - Basic machining operation and important parameters (KALPAKJIAN; SCHMID,
2000).
Cutting tools that can be used in machining are classified into three groups:
1. Single-Point Tools: these are cutting tools that have one cutting part and one shank
as illustrated in Figures 2.10 (a). Although these tools have traditionally been
produced from solid tool-steel bars, they have been largely replaced by carbide or
other indexable inserts of various shapes and sizes, as shown in Figure 2.10 (b).
Turning operation is probably the most common of all the machining processes. It is
26
defined as the process for generating external surfaces (BOOTHROYD; KNIGHT
(1989), KALPAKJIAN; SCHMID (2000)). The workpiece material in the form of a
cylindrical bar is held in the chuck of a lathe and rotated about the axis of symmetry
while the cutting tool is held in a tool post and moved with a feed motion parallel to
the workpiece axis (Figure 2.11). Single point tools can also be used in other cutting
operations such as shaping, boring, trepanning and threading (BOOTHROYD;
KNIGHT, 1989).
Figure 2.10 - Schematic illustration of typical single-point cutting tool with the tool angles
(KALPAKJIAN; SCHMID, 2000).
Tool
(insert)
Toolholder
Workpiece
Tailstock
Cross slide
Center
Toolpost
Chuck
Figure 2.11 - Workpiece-tool-machine system for turning operation.
27
2. Multi-Point Tools: are cutting tools that have a more than one cutting part and they
are secured to a common body such as drills, milling cutters, reamers, etc. The
majority of multipoint tools are intended to be rotated and have either a taper or
parallel shank for holding purposes or a bore through which a spindle can be
inserted. Milling is the most common operation using muilti-points tools. Unlike
turning operation, with the standard method of milling the tool rotates on its own
axis, in a fixed position, and the workpiece is brought into contact with this rotating
tool. The workpiece is then moved in the required direction to carry out the
machining process (Figure 2.12). Other cutting operations that use multi-point tools
are drilling, reaming, tapping, broaching and sawing (BOOTHROYD; KNIGHT,
1989);
Chuck
Tools
(inserts)
Milling
cutters
Workpiece
Figure 2.12 - Form-milling operation with gangs of side and face milling cutters (AB
SANDVIK COROMANT, 1994).
3. Abrasive Wheels: these are cutting tools with a large number of small cutting parts
that are randomly shaped and oriented. The machining process that uses abrasive
wheels as a cutting tool is termed abrasive machining. The most common abrasive
machining process is grinding where the abrasives are bonded to the shape of a
wheel (or grinding wheels) which rotates at a high speed. Grinding wheels consist
of individual grains (each grain is considered as small cutting part) of very hard
28
material (usually silicon carbide or aluminum oxide) bonded in the form required.
Typical shape of abrasive wheels are cylindrical, disc-shaped, or cup shaped. The
other finishing operations using abrasives are honing and lapping (GHOSH;
MALLIK, 1986).
2.4.2 Nomenclature of Cutting Tools
In this Section more emphasis will be placed on nomenclature of single-point cutting
tools. Figures 2.9 and 2.10 (a) show the general terms used for single-point cutting tools.
The part of the tool engaged in the cutting process is characterised by the rake and flank
faces.
• Rake Face: the surface or face of the tool over which the chip flows during the
machining process;
• Flank Face: the surface of the tool in contact with the freshly cut surface of the
workpiece. The part of the flank which intersects the rake face to form the cutting
edge is termed major flank. The remaining part of the flank, which is not actually
engaged in cutting and intersects the rake face, is called the minor flank or minor
clearance face (See Figure 2.10 (a)).
In order to prevent interference and rubbing of the freshly cut surfaces against the tool,
clearance angles are provided on both the flank faces. The one on the major flank face is
called the clearance angle and the one on the minor flank face is called the end clearance
angle. In order to strengthen the cutting wedge a radius is always provided at the point where
both the flank faces meet. This is called the nose radius.
System of planes are generally used to better define and specify the angles of a cutting
tool. If one point on the cutting edge is selected and a reference system of planes is located in
this point. The two major systems of planes are known as “tool-in-hand-system” and “tool-in-
use-system” (Figure 2.13) (BOOTHROYD; KNIGHT, 1989). The former is used to define the
geometry of the tool for its manufacture and measurement, while the later is used to specify
the geometry of the cutting tool when it is performing a cutting operation. Other detailed
definitions can be found elsewhere (BOOTHROYD; KNIGHT (1989), BRITISH
STANDARD (1972)).
29
(a)
(b)
Pn: cutting edge normal plane Ps: working cutting edge plane
Ps: tool cutting edge plane Pr: working reference plane
Pr: tool reference plane
Figure 2.13 - Cutting tool planes: (a) “tool-in-hand” planes and (b) “tool-in-use” planes
(BOOTHROYD; KNIGHT, 1989).
Angles derived from “tool-in-hand-system” are called “tool angles” and angles derived
from the “tool-in-use-system” are known as “working angles”. These angles are showed in
Figure 2.14 and listed below:
• Tool cutting edge angle (k
r
);
• Tool minor cutting edge angle (k´
r
);
• Tool included angle (
ε
r
);
• Tool cutting edge inclination angle (
λ
s
);
• Tool normal rake angle (γ
n
);
• Tool normal clearance angle (
α
n
);
• Tool normal wedge angle (
β
n
).
In the suffix after the symbol of the angles of “tool-in-hand” planes (Figure 2.13)
represents the plane of the system in which the angle is measured. The suffix ‘r’ refers to the
plane Pr (tool reference plane) that is parallel to the tool base. The suffix ‘s’ refers to the plane
Ps (tool cutting edge plane), which is tangential to the cutting edge and perpendicular to the Pr
plane. The suffix ‘n’ refers to the plane Pn (cutting edge normal plane), which is
perpendicular to the cutting edge. For the “tool-in-use” planes the suffix ‘re’ refers to the
30
plane Pre (working reference plane) which is perpendicular to the resultant cutting direction;
the suffix ‘se’ refers to the plane Pse (working cutting edge plane) which is tangential to the
cutting edge and perpendicular to the working reference plane Pre and the suffix ‘n’ refers to
the plane Pn (cutting edge normal plane), which is perpendicular to the cutting edge.
Figure 2.14 - Tool angles for a single-point tool according to the ISO: tool cutting edge angle
(k
r
), tool minor cutting edge angle (k´
r
), tool included angle (
ε
r
), tool cutting edge inclination
angle (
λ
s
), tool normal rake angle (γ
n
), tool normal clearance angle (
α
n
) and tool normal
wedge angle (
β
n
) (BOOTHROYD; KNIGHT, 1989).
.
2.5 Chip Formation Process
The understanding of the metal cutting process is much about the behaviour of various
types of metals during deformation to produce chips. Many parameters in metal cutting, such
as tool wear, friction between the tool rake face and the chip, cutting forces, temperature,
surface finish, machined surface integrity and machining power are related to the process of
chip formation. In this context, understanding the mechanisms and process of chip formation
is essential as the fundamental knowledge can help in resolving practical problems (XIE;
BAYOUMI; ZBIB, 1996) as well as in optimising cutting conditions to increase the
production rate which will further reduce the production costs. The type of chips produced
31
during metal cutting depends on the material being machined and the cutting conditions
employed.
Metal cutting is a process of removing a thin layer from the surface of a workpiece
material by driving a wedge shaped tool asymmetrically into it. During this process two new
surfaces are generated and the thin layer of material, termed chip or swarf undergoes
deformation throughout its volume. The chip impinges on the tool rake face while moving in
a direction away from the direction of motion of the workpiece. The chip then separates from
the workpiece via shear deformation occurring along the shear plane.
The chip formation can be explained considering the “klmn” volume of metal travelling
against the direction of the tool edge (Figure 2.15). As this volume of material passes through
the primary shear zone (boundary between the unsheared work material and the body of the
chip), it deforms plastically to a new shape “pqrs”. As shearing continues, the volume (chip)
slides across the rake face of the tool and then it is forced to change direction, experiencing a
curvature and eventually curling. When the strain in the chip reaches a critical value, the chip
fractures and the process begins over again (SHAW, 1984). This region, the interface between
the tool and the chip on the rake face, is termed secondary shear zone. The chip is thicker than
the actual depth of layer to be removed (underformed chip thickness or depth of cut, t
1
), i.e., t
2
(actual chip thickness) > t
1
and the chip is correspondingly shorter. Also, the chip velocity
(V
c
) is, in the same proportion, smaller than the cutting speed V, i.e, V > V
c
. With the purpose
of simplification, the primary shear zone is assumed to be a plane, as illustrated in Figure 2.16
(THE METALS HANDBOOK, 1989). The shear angle, denoted by
φ
(schematically
illustrated in Figures 2.15 and 2.16), is the angle between the shear plane and the direction of
cutting and can be determined by direct measurement of t
1
and t
2
given by the Equation (2.4):
α
α
φ
sin
r
r
×−
×
=
1
cos
tan
(2.4)
Where r (the chip thickness ratio) = t
1
/t
2
, with r << 1, and
α
: tool rake angle.
32
Figure 2.15 - Metal cutting diagram - the chip formation (TRENT; WRIGHT, 2000).
Figure 2.16 - Metal cutting diagram illustrating the primary and secondary shear zones (THE
METALS HANDBOOK, 1989).
33
The shear strain within the primary shear zone is influenced by the value of the shear
plane angle and has a significant effect on the chip form. Small shear plane angle means high
strain and vice-versa. If the chip thickness ratio is big, the shear plane angle is small and the
chip moves away slowly (e.g. for aluminium), while a large shear plane means a thin, high-
velocity chip (e.g. for steel) (TRENT; WRIGHT, 2000). A higher strain rate favours the
formation of discontinuous chips while lower strain rate produce continuous chips as the shear
thickness of element (“klmn” volume of metal) is bigger. Because higher shear strain rate
occurs at very short time interval, which does not allow the metal to recover from the strain,
chip fractures into discontinuous chips. It is known that high energy is consumed in the
secondary shear zone as a consequence of a resistance to chip motion. This resistance to chip
motion reduces the shear plane angle and extends the length of the primary shear zone and
vice-versa (WRIGHT; BAGHI; CHOW, 1982). The shear plane angle,
φ
, and thus the chip
thickness ratio, r, therefore depend on the work material, the tool material, the tool geometry
(mainly of tool rake face) and the cutting conditions. In order to determine the value of the
shear plane angle (
φ
), Ernst & Merchant (1940) assumed the position of the shear plane would
be take up such a value as to reduce the work done in cutting to a minimum, or in order words
a value such that the shear stress acting upon the primary shear plane would be a maximum.
The resulting expression is given by the Equation (2.5) (BOOTHROYD; KNIGHT, 1989):
()
karcot2 =−+
α
ρ
φ
(2.5)
Where
φ
= shear plane angle;
α
= tool rake angle; k = constant for the work material,
which is related to change of the shear strength with the compressive strength and therefore
changes from material to material.
ρ
= angle of chip friction, which can be calculated by the Equation (2.6):
()
N
F
==
ρµ
tan
(2.6)
Where
µ
is the coefficient of friction, F is the force required to initiate or continue
sliding and N is the force normal to the interface at which sliding takes place.
2.6 Classes of Chips
As commented in Section 2.5, the type of chip produced during metal cutting is mainly
associated to the properties of the material being machined and the cutting conditions
34
employed. Based on these parameters, chips can be basically grouped into four classes
(TRENT; WRIGHT (2000), BOOTHROYD; KNIGHT (1989)):
a) Continuous chips: These are usually formed when cutting ductile metals and alloys
that do not fracture on the shear plane (TRENT; WRIGHT, 2000) such as
aluminium, low alloys steels and copper. This chip can also be generated when
machining single phase materials or multi-phase materials at high cutting speed and
with tools that have larger rake angles (positive rake angles). Figure 2.17 (a) shows
the schematic representation of this chip type. In the process of cutting, highly
distorted crystals in the chip are sheared off from the parent metal in the shear zone
with very large amounts of shear strain, remaining in a homogeneous form and not
fragmenting. Continuous chips are undesirable because they usually wrap
themselves around the workpiece or get tangled around the tool holder, thus
adversely affecting the surface finish generated and/or causing tool damage. In
some cases machining has to be interrupted in order to clear them away. Cutting
tools with chip breakers can be used to intermittently break the long continuous
chips into pieces so that machining does not need to be interrupted. Continuous
chips may adopt many shapes – straight, tangled or with different types of helix;
b) Continuous chips with built-up-edge (BUE): They are a variation of the continuous
chips observed at the chip-tool interface when cutting a certain group of materials
(those with two phases in their structures) at lower cutting speeds (Figure 2.17 (b)).
BUE formation involves work hardening and crack initiation/propagation when
particles of the workpiece material weld to the rake face of the tool during
machining. The BUE is a dynamic structure that consists of successive layers of
material from workpiece which are gradually deposited on the tool rake face (tip of
the tool), Figure 2.15 (b). As the deposition becomes larger it displaces the chip
from direct contact with the tool. When the size of the BUE reaches a certain value,
at which the resolved shear stress is high enough to shift the shear zone into the
BUE, parts of its structure is carried away on the work surface or on the underside
of the chip. Detailed information on BUE formation is available elsewhere in
(TRENT; WRIGHT (2000), MACHADO; WALLBANK (1990)). The occurrence
of BUE is an important factor that affects surface finish and tool wear;
35
(a)
(b)
(c)
Workpiece
Chip
Built-up edge
Tool
Workpiece
Tool
Chip
Workpiece
Chip
Tool
(d)
Figure 2.17 - Classes of chips: (a) Continuous chip, (b) Continuous chip with BUE,
(c) Discontinuous chip, (d) Serrated chips (TRENT; WRIGHT (2000), MACHADO;
WALLBANK (1990), KALPAKJIAN; SCHMID (2000)).
36
c) Discontinuous chips: These are usually produced when machining brittle materials
containing inclusions and impurities at low or high cutting speeds with large depth
of cut and using tools with small rake angles. Impurities and hard particles increase
the possibility of crack propagation in the material while bigger depth of cut
increases the chances of cutting through such defective regions resulting in
generation of discontinuous chips. The chips are produced when the work material
is not able to sustain large amounts of severe strain without fracture (Figure 2.17
(c)). With increasing cutting speed the chip tends to become more continuous
because it is more difficult for contaminants to penetrate at the interface and reduce
the constraining force;
d) Serrated chips: Also designated as segmented chips or “saw toothed chips” and
described as chips with variation in thickness. The strain involved here is confined
to narrow bands between segments with very little deformation within these
segments. Adiabatic shear may occur during chip formation, particularly when
machining materials exhibiting poor thermal conductivity and high strength that
decreases abruptly with temperature such as titanium alloys. In this case serrated
chip forms due to unstable adiabatic shear. Komanduri and Brown (1981) proposed
that a mechanism of instability in the primary shear zone is based on the behaviour
of certain polycrystalline metals at large strains under combined shear and
compression, exhibiting negative shear stress-shear strain characteristics and
involving microcracking. Serrated chips have saw tooth like appearance as
illustrated in Figure 2.17 (d). Tool wear is adversely affected by the formation of
serrated chips.
2.7 Forces in Metal Cutting
The understanding of the forces acting on the cutting tool and the study of their
behaviour are of vital importance in the design and manufacture of machine tools and their
components as well as in optimising tool design and controlling the surface finish and surface
integrity of machined components. The cutting forces also act as a machinability index.
37
The cutting force can be defined as the force exerted by the tool cutting edge on the
workpiece material to promote chip shearing. The relative motion between the workpiece
material and the cutting tool is responsible for generation of this force. Component forces in a
machining process can be resolved into three components in an orthogonal reference system
usually involving the machine tool coordinate system (Figure 2.18 (a)). These three
components forces are represented by cutting force (F
c
), feed force (F
f
) and the radial force
(F
r
) (also called the thrust force). The cutting force is usually the largest force acting on the
tool rake face in direction of the cutting velocity. Feed force acts parallel to the direction of
the tool feed. The radial force acts to push the tool away from the workpiece in the radial
direction. This force is usually the smallest of the force components in semi-orthogonal
cutting and, for purposes of analysis of cutting forces in simple turning, it is usually ignored
(DE GARMO; BLACK; KOHSER, 1999). Forces acting in metal cutting are generally
measured with dynamometer which is a force-platform device that incorporates piezo-electric
load cells.
The magnitude of component forces during a machining operation are mainly dependent
to the work material and the cutting speed, feed rate, depth of cut, tool geometry, tool
material, wear conditions of the tool and cutting fluids employed. The maximum compressive
stress acts on the cutting edge and reduces to zero when approaching the end of the tool-chip
contact zone. Turning materials which produce discontinuous chips tend to produce lower
forces due to the shorter tool-chip contact length. During machining of materials with high
strength, large forces are required due to the higher stresses on the shear plane which
subsequently result in large amount of heat generated in the cutting zone. With the increase in
heat the yield strength of the tool material falls rapidly (TRENT; WRIGHT, 2000). The large
forces combined with high cutting temperature will also accelerate wear at the cutting edge
during machining. Another factor which has important influences in cutting forces is the
cutting speed. An increase in cutting speed increases the shear plane angle, this helps to
generate thinner chips resulting in reduced tool-chip contact length, thereby causing a drop in
cutting forces (JAWAID, 1982). Increase in cutting speed also increases the cutting
temperature (TRENT; WRIGHT, 2000). The rise in temperature promotes the softening of
work material, and thus a reduction in the shear strength on the shear plane. Reduction in
shear strength leads to a decrease in the component forces. Component forces increases as the
tool is worn out because the area of contact at the clearance face increases as the wear at the
flank face progresses.
38
(a)
(b)
Figure 2.18 - Cutting forces a) Three components forces acting on the cutting tool (DE
GARMO; BLACK; KOHSER, 1999) and b) Merchant´s circle (TRENT; WRIGHT, 2000).
Cutting tool geometry, especially the rake angle, also affects the component forces
during machining. Tools with larger rake angles require lower cutting and feed forces and
reduces the strength of the tool edge due to reduction of the included angle. Tools with
negative rake angles are stronger because the higher included angle. Cutting fluids/lubricants
can also influence the cutting forces acting to prevent seizure between the tool and workpiece
material at lower speed conditions, thus reducing forces acting on the tool.
39
2.8 Stress and Strain Distribution in Machining
2.8.1 Stress Distribution
Determination of forces also helps to work out the stresses encountered by the tool
during the cutting process. Assuming a uniform material distribution on the shear plane and at
the tool-chip interface, forces can be resolved into normal and shear forces (stress) and
calculated using the following expressions, Equations (2.7) and (2.8) (TRENT; WRIGHT,
2000):
Normal stress:
S
N
A
F
=
σ
(2.7)
Shear stress:
S
S
A
F
=
τ
(2.8)
Where F
N
and F
S
are the normal and tangential forces acting on the shear plane
respectively (Figure 2.18 (b)).
A
S
is the area of the shear plane. In orthogonal cutting, it is calculate by the Equation
(2.9):
φ
sin
DOCf
A
S
×
=
(2.9)
Where f is the undeformed chip thickness (feed), DOC is the depth of cut and
φ
= shear
plane angle.
Therefore, the force required to form the chip depends on the shear yield strength of the
work material under the cutting conditions and the area of the shear plane. The normal stress
on the shear plane has no effect on the magnitude of the shear stress and decreases with
increasing rake angle. The shear stress on the shear plane is independent of the rake face. The
forces increase in direct proportion to increments in the feed rate and depth of cut, the two
major variables under the control of the machine tool operator. For a zero-degree rake angle
tool, the shear plane may vary from a maximum of approximately 45º to a minimum of 5º or
even less (TRENT; WRIGHT, 2000).
According to Zorev (1963), on the rake face the compressive stress has a parabolic
distribution, being zero where the chip loses contact with the tool and having a maximum
value at the cutting edge, (Figure 2.19) and can be calculated by the Equation 2.10)
(BOOTHROYD; KNIGHT, 1989):
40
y
c
xq×=
σ
(2.10)
Where x is the distance from the point where the chip loses contact with the tool, q and
y are constants.
Figure 2.19 - The Zorev´s model of stress distribution on the rake face of a cutting tool in
orthogonal cutting where σ
fmax
= maximum normal stress, σ
f
= normal stress, τ
f
= shear stress,
τ
st
= shear strength of chip material in the sticking region (BOOTHROYD; KNIGHT, 1989).
Analysis of the stress distribution also showed that the shear stress is constant at the
sticking region and equal to the shear strength of the work material. This stress drops to zero
at the sliding region where the chip loses contact with the tool. When the stress is very high
the tool is deformed downwards towards the clearance face. This form of deformation
increases tool forces, and thus accelerating wear processes and consequently reducing tool
life. Stress distribution also varies with the work material due to the frictional coefficient
between chip and tool and the deformation characteristics of the work material (KATO;
YAMAGUSHI; YAMADA (1972), CHILDS; MAHDI (1989)). As the cutting speed is raised,
the temperature rises and the yield stress of the tool reduces. These tend to weaken the tool
material resulting deformation.
2.8.2 Strain Distribution
Strain in machining is generally related to the shear strain at the primary shear plane
(Figure 2.20) and calculated by the expression (Equation 2.11):
41
)cos()(
)cos(
γφφ
γ
δ
−
=
∆
∆
=
xsinY
S
(2.11)
Figure 2.20 - The shear strain in the shear plane (SHAW, 1984).
When the chip thickness ratio is near to unity, i.e., the chip is thin, the shear strain is
usually around 2. As this factor increases shear strain value may reach 5 or even greater
(TRENT, 1988a). In the particular case of a serrated chip with adiabatic thermoplastic shear,
the amount of strain in the flow zone is much higher than the strain in the primary shear plane.
The ability of metals and their alloys to withstand such strains in the flow zone (zone of
extremely intense shear formed on chip-tool interface) without fracture is attributed to the
very compressive stresses in this region which inhibit the initiation of cracks, and cause the
re-welding of such small cracks as may be started or already existed in the work material
before machining (TRENT; WRIGHT, 2000). Serrated chips are mainly generated when
machining titanium and its alloys. When machining such material the shear strain is around 8
within the shear bands and 1.3 within the segments, on the basis of metallographic
observations (TURLEY; DOYLE; RAMALINGAM, 1982). Furthermore, the strain rate in the
shear plane is up to 1000 s
-1
, which is considered high (SHAW, 1984). Figure 2.17 (d)
illustrates an adiabatic shear band.
A model of shear strain within the flow zone in the primary shear plane was proposed
by Trent and Wright (2000). In this model the amount of shear strain increases with the
distance from the body of the chip-flow zone interface to the tool rake face. Theoretically, the
amount of shear strain becomes infinite at the tool face but laminar flow cannot be considered
42
as persisting closer to the tool surface than a few microns due to the surface roughness of the
tools.
2.9 Heat Generation During Machining Operation
When a material is deformed elastically, some energy is spent to increase its strain energy,
which is returned during unloading. For plastic deformation this is different. Here most of the
energy used for shearing the workpiece is converted into heat. In a machining operation
practically all of the energy involved is consumed either in plastic deformation or in friction.
These essentially end up as thermal energy, heat (GHOSH; MALLIK, 1986). A small portion
of the remaining energy is retained in the system as elastic energy while the rest is used for
generation of new surface area.
There are three zones where the heat is generated during machining as shown in Figure
2.21 and listed below:
i) The primary shear zone, where significant deformation occurs. A large portion of
the heat generated in this zone goes to the chip and the rest is conducted into the
workpiece and raises its temperature, some times causing dimensional alteration of
the workpiece;
ii) The secondary shear zone (the tool-chip interface). The heat generated in this zone
is due to the sliding motion of the chip on the rake surface of the tool and is
responsible for the rise in tool temperature. In this zone the chip also takes away the
major portion of the heat.
iii) The workpiece-clearance face interface zone. The heat generated in this zone is due
to friction of the work material against the flank face of the tool. This zone is
considered less important, unless a very small clearance angle is used or the tool is
severely worn. When sharp tools are employed, the contribution of this heat source
zone is negligible GHOSH; MALLIK, 1986).
43
(a)
(b)
Figure 2.21 - Zones of heat generation during machining: (a) schematic diagram,
(b) isothermal lines for dry orthogonal cutting of free machining steel with carbide tool
(α = 20º) obtained from a finite element technique, at a cutting speed of 155.4 m min
-1
and a
feed rate of 0.274 mm rev
-1
[adapted from SHAW (1984)]
2.9.1 Effect of Cutting Parameters on temperature generated during machining
Temperature influences the cutting action in several important ways, including
(BEDDOES; BIBBY, 1999):
• Altering the properties of the machined surface;
• Causing dimensional changes to the work material and adversely affecting
dimensional accuracy of machined component;
• Adversely affecting the strength, hardness and wear resistance of the cutting tool.
C
Work
p
iece
A
Primary
shear zone
Secondary
shear zone
B
D
44
The strength and thermal conductivity of the work and tool materials influence the
maximum temperature during machining. Increasing feed rate, cutting speed and depth of cut
all influence also have significant effect on the cutting temperature, according to the relation
given by the Equation (2.12):
T
M
α
V
a
×
f
b
×
DOC
c
(2.12)
Where T
M
is the mean cutting temperature, V is the cutting speed, f is the feed rate,
DOC is the depth of cut and a, b and c are constants (a > b > c).
Reduction in hardness and wear resistance of the tool with increasing temperature is the major
factor that adversely affects tool life. An increase in the metal removal rate leads to a
proportional increase in temperature on the cutting tool. This has a direct influence on tool
wear and tool life. The tool temperature may not be a critical problem during machining of
low strength and low melting point materials such as aluminium and magnesium. However,
when ferrous and other high strength materials such as cast iron, steel, nickel and titanium
alloys are machined, temperature rises with the speed and the tool strength decreases, leading
to a faster wear and consequent tool failure (TRENT; WRIGHT, 2000). Although machining
at a high speed is desirable, for higher productivity, the faster tool wear due to the high
temperature generated at the cutting zone will establish a limit to the practical cutting speed
for each tool-work material pair. Different work materials, during machining, generate
different amounts and patterns of heat. For example, the point of maximum temperature when
machining titanium alloys is located at the tool rake face and on the flank face when
machining nickel alloys (JAWAID, 1982).
When machining materials with high strength, large forces are required due to the
higher stresses on the shear plane, leading to large amount of heat generation in the cutting
zone. The large forces combine with high cutting temperatures generated will also accelerate
failure of the cutting edge during machining An increase in cutting speed increases the cutting
temperature (GHOSH; MALLIK (1986), BOOTHROYD; KNIGHT (1989)), consequently
encouraging softening of the work material (TRENT; WRIGHT, 2000).
Some considerations regarding metallurgical parameters influencing temperature during
machining were enumerated by Trent (1988b):
• The melting point of the main element of the work material. The higher the melting
point the higher will be the temperature at the tool-work material interface for any
cutting speed;
45
• Alloying elements which strengthen a metal promotes increase in temperature at the
tool-work material interface at any metal removal rate;
• Introduction of phases into the work material that easily form sheared interface
layers. They act reducing temperature.
2.9.2 Heat Generation and cutting temperature when machining titanium alloy
The understanding of heat generation process and temperature behaviour in machining of
titanium alloys plays an important role in distribution of cutting tool temperature. The high
temperatures generated close to the cutting edge of the tool are the principal reasons for the
rapid tool wear commonly observed (TRENT; WRIGHT, 2000). In comparison with steel, the
heat capacity of titanium and its alloys is much reduced. The consequence is that a
considerably greater portion of heat that is generated during cutting enters into the tool
because it cannot be removed with the fast flowing chip or bed into the work material due to
the low thermal conductivity of titanium alloys, about 37% and 86% lower than the thermal
conductivity of Inconel 718 alloy and AISI 1045 steel, respectively (MANTLE;
ASPINWALL, 1998). About 80% and 50% of the heat generated is absorbed in the tool when
machining titanium-base, Ti-6Al-4V, alloy and Ck 45 (AISI 1045) steel respectively
(KONIG, 1979), as illustrated in Figure 2.22. When machining a titanium alloy at a cutting
speed of about 30 m min
-1
the temperature developed at the cutting edge of a carbide tool is
about 704ºC, while for steel the temperature is about 538ºC (ZLATIN; FIELD, 1973). The
influence of cutting speed on temperature generated when machining commercially pure
titanium and various titanium alloys with cemented carbide (K10 grade) was investigated by
Motonishi et al. (1987) (Figure 2.23). Increase in cutting speed resulted in higher cutting
temperature. The cutting temperature of Ti-6Al-4V alloy was about 200ºC higher than pure
titanium, S45C.
46
0
20
40
60
80
100
0 42 84 126 168
Ratio of heat flow
Chi
p
Tool
Ti-6Al-4V
%
Steel Ck 45
Carbide K10, K20
Oxide ceramic
Carbide P10
Diamond
High speed steel
Stelit
Thermal Conductivity
λ (
J/mm s
°
C)
Figure 2.22 - Distribution of thermal load when machining titanium-base, Ti-6Al-4V and Ck
45 (AISI 1045) steel [adapted from KONIG (1979)].
400
600
800
1000
1200
50 100 150 200
Cutting speed (m/min)
Cutting temperature (ºC)
Ti-5Al-2.5Sn
Ti-6Al-4V
P10-S45C
C.P. Ti (KS50)
Figure 2.23 - Influence of cutting speed on the cutting temperature when machining titanium
and its alloys [adapted from MOTONISHI et al. (1987)].
47
2.10 Tool Failure Modes
During machining, as the work material is cut, the tool edge degrades away causing a
gradually alteration of tool shape, consequently affecting dimension, tolerance and the quality
of machined part and associated reduction of the efficiency of the machining process. In metal
cutting, progressive wear occurs on both the rake and the clearance faces of cutting tools. The
type of tool wear depends on tool geometry, tool and work material and their physical,
mechanical and chemical properties as well as cutting parameters and machining environment
employed. Figure 2.24 shows diagrammatically typical tool failure modes (DEARNLEY;
TRENT, 1985):
a) Flank wear
b) Rake face wear (crater wear)
c) Notch wear
d) Cutting edge chipping
A: Flank wear; B: Rake face wear (crater wear); C: Notch wear; D: Cutting edge chipping
Figure 2.24 - Regions of wear on a cutting tool (DEARNLEY; TRENT, 1985).
a) Flank wear occurs at the flank face of the tool as the newly formed surface rubs
against of the clearance face of the tool causing adhesive and/or abrasive wear which
is enhanced by the rise in temperature during machining. The friction action causes
the loss of relief angle on the clearance face of the tool, thereby resulting in wear. In
general, the flank wear rate is high at the beginning of the cut, but decreases as the
wear land (VB) reaches a critical value. As the cutting time progresses a second
48
critical value is reached. Further machining results in a rapid increase in flank wear
rate. Machado (1990) observed that maximum flank wear (VB
max
) was the
predominant form of wear when machining titanium-base, Ti-6Al-4V alloy, with
uncoated cemented carbides under conventional and high pressure coolant supplies.
It was also observed that a layer of work material was always bonded to the tool, on
the wear land and can some times be thicker, especially when machining with
conventional coolant flow. Motonishi et al. (1987) reported that flank wear on
carbide tools after machining titanium alloys is associated with fluctuations of
cutting force caused by the serrated chips;
b) Rake face wear (crater) is generally associated with high temperatures generated at
the chip-tool interface and occurs on the rake face of a tool due to a combination of
diffusion and adhesion as the chip moves over the rake face of the tool. The
maximum depth of the crater usually occurs near the midpoint of the contact length
between the chip and the rake face, where the temperature is believed to be at a
maximum. Excessive crater wear alters the geometry of the cutting edge and can
adversely affect chip formation and weaken the tool because of the decrease in its
yield strength (AB SANDVIK COROMANT, 1994). It was reported that during
machining of titanium-base, Ti-6Al-4V alloy, with cemented carbides a shallow
crater was initially formed on the rake face of the tool and a little flank wear was
produced along the whole extension of the depth of cut (MACHADO, 1990).
Additionally, the crater was formed close to the cutting edge and joining the flank
wear;
c) Notch wear or groove formation occurs as a result of sliding wear when machining
work materials which have their surfaces strained hardened from previous cuts or
work hardened due to coolant effects and/or fluctuating temperature. Leading edge
notching is a major problem when machining titanium and aluminium alloys
(hardening by precipitation of γ’ phase) (RICHARDS; ASPINWALL, 1989). Stress
concentration towards the unmachined surface due to the stress gradient and the
presence of a burr at the edge of the freshly machined surface as well as the cut
surface also contribute to the notch formation. Jawaid (1982) suggested that
notching generally takes place under sliding conditions and disappears at cutting
conditions where seizure takes place. It was also reported that the presence of
different machining environments, such as gases like air, argon and nitrogen,
49
influence notch wear rate when machining nickel-base alloys with ceramic tool
materials. Notching at the depth of cut was the predominant form of wear observed
when machining nickel-base, Inconel 901 alloy, with various grades of carbides
(MACHADO, 1990);
d) Cutting edge chipping is the process of plucking of small pieces of tool particles at
the cutting edge of the tool. The chipped off particles can be small or large
fragments. Unlike wear, chipping is not a gradual process. It occurs on a random and
unpredictable manner and alters the geometry of the tool edge. Chipping is
associated with mechanical shock due to impact, particularly in interrupted cutting.
Excessive wear for the tool such as flank wear, crater wear and notching will
increase cutting forces and weaken the tool. The presence of thermal cracks
generated in interrupted cutting due to the thermal cycling can also lead to the
generation of chipping and sudden failure of the cutting tool. High feed rates and
depth of cut can also cause chipping. Chipping was observed when machining with
tool materials with inadequate fracture toughness such as ceramic tools and finishing
grades of tungsten carbide containing less than 3% weight of cobalt (SHAW, 1984).
2.11 Tool Wear Mechanisms
The understanding of the wear mechanisms of tools in metal cutting has great importance for
improvement and development of better tool materials and designs in order to minimise tool
wear. Detailed research into tool failure modes has suggested that wear in metal cutting is
attributed to the following wear mechanisms schematically illustrated in Figure 2.25 and
listed below (TRENT; WRIGHT, 2000):
a) Attrition wear
b) Abrasion wear
c) Diffusion wear
d) Plastic deformation
50
Workpiece
Workpiece
Workpiece
Workpiece
Figure 2.25 - The main wear mechanisms on a cutting tool [adapted from (TRENT;
WRIGHT, 2000)].
a) Attrition wear is also known as adhesive wear and is a mechanism that occurs at
relatively low cutting speeds which changes the geometry of the tool by mechanical
detachment of small particles from tool surface which are carried away by the work
material. Attrition wear is more related to the irregular flow of the chip as it passes
the cutting edge and the ability of the tool to withstand the tearing action resulting
from uneven flow under conditions of sliding. The high compressive stresses acting
at the tool-workpiece interface causes adhesion at the asperity. Fragments of the
work material are torn and carried along the tool edge. As the rubbing action of the
tool against the workpiece continues the welded material breaks away carrying with
it particles of the tool material. This is a process that involves random removal of
whole grain or small fragments of the tool material resulting in rapid tool wear. The
aspect of worn tool surface is uneven and nibbled away (TRENT; WRIGHT, 2000).
The factors that influence and promote attrition wear are low speeds and feed rates,
lack of rigidity of the machine tool, vibration and the formation of built-up-edge;
b) Abrasion wear involves the loss of tool material by a hard particles trapped between
the tool and the work material. The particles could be dislodged tool materials,
fragments of built-up-edge or hard carbides and oxides already existing in the
workpiece. The high compressive stresses acting at the tool-workpiece interface
51
keep trapped hard particles between the machined surface and the tool. As the tool
rubs over the machined surface, the particles plough through the tool removing
material at the flank face. Abrasive wear at the rake face of the tool is due to sliding
action of hard particles located at the under side of the chip as it passes over the
rake face of the tool;
c) Diffusion wear involves exchange of atoms between the tool and work material at
the tool-workpiece and the tool-chip interfaces. It is a mechanism that occurs at
elevated temperatures. Other factors that influence the rate of diffusion tool wear
are metallurgical relationship between the work material and the tool material
(solubility). The feature characteristic of diffusion wear is the smoothness of the
worn surface;
d) Plastic deformation occurs as a result of gradual crack growth caused by thermal
shocks and compressive stresses acting on the rake face of the tool. Since there is no
lost of tool material during plastic deformation, it is not considered a physical wear
process. This process is more likely to occur at high feed rates or when machining
materials with poor machinability characteristics like superalloys and certain steels
that generate high temperatures and compressive stress which may lead to sudden
fracture. Increase in localised stress and temperature can accelerate wear on the
flank face or tool nose leading to the subsequent fracture.
2.12 Titanium Machinability
The initial studies on the machining of titanium alloys, the tool wear guidelines and the
chip formation process, started around the 1950s in the United States and in France (COOK
(1953), COLWELL; TRUCKENMILLER (1953), SIEKMANN (1955)). Titanium alloys are
attractive material to designers in the aerospace industry because of their unique combination
of strength and lightness despite their low thermal conductivity that concentrates heat
generated during cutting at the cutting edge (YANG, 1970). They also exhibit high chemical
reactivity, tendency to form adiabatic shears bands and/or rapid work hardening during
machining (KOMANDURI; BROWN (1981), TURLEY; DOYLE; RAMALINGAM (1982),
SHAW (1967), CHILD; DALTON (1968), KOMANDURI; VON TURKOVICH (1981)).
These pose considerable problems in manufacturing hence titanium alloys have poor
52
machinability (FIELD (1968), EZUGWU; BONNEY; YAMANE (2003), EZUGWU; WANG
(1997), HONG; MARKUS; JEONG (2001), MACHADO; WALLBANK (1990),
DEARNLEY; GREARSON (1986), ZLATIN; FIELD (1973), MOTONISHI et al. (1987),
DEARNLEY; TRENT (1985), MACHADO (1990), COOK (1953), COLWELL;
TRUCKENMILLER (1953), SIEKMANN (1955), YANG (1970), SHAW (1967), CHILD;
DALTON (1968), KOMANDURI; LEE (1984), JAWAID; CHE-HARON; ABDULLAH
(1999), HONG; DING; JEONG (2001), EZUGWU; OLAJIRE; WANG (2002),
VENUGOPAL et al. (2003)) and are referred to as difficult-to-machine materials.
Machinability can be described as how easily a material can be cut to the desire shape
(surface finish and tolerance) with respect to the tooling and machining processes involved.
Efforts to improve the high temperature properties of titanium alloys usually impair their
machinability (BOYER, 1996). Other factors and output variable such as tool life, quality of
machined surfaces (surface texture, surface integrity and form tolerances), chip shape,
component forces and power consumed during the machining operation are the main
parameters that determine the machinability of a material. Reasons for the relatively poor
machinability of titanium alloys have been discussed elsewhere and summarised below
(FIELD (1968), SECO TOOLS (2002a), EZUGWU; BONNEY; YAMANE (2003),
EZUGWU; BONNEY; YAMANE (2003), EZUGWU; WANG (1997), HONG; MARKUS;
JEONG (2001), LOVATT; SHERCLIFF (2002), MACHADO; WALLBANK (1990),
DEARNLEY; GREARSON (1986), ZLATIN; FIELD (1973), MOTONISHI et al. (1987),
DEARNLEY; TRENT (1985), MACHADO (1990), COOK (1953), COLWELL;
TRUCKENMILLER (1953), SIEKMANN (1955), YANG (1970), SHAW (1967), CHILD;
DALTON (1968), KOMANDURI; LEE (1984), JAWAID; CHE-HARON; ABDULLAH
(1999), HONG; DING; JEONG (2001), EZUGWU; OLAJIRE; WANG (2002),
VENUGOPAL et al. (2003), OLOFSON (1965), CATT; MILWAIN (1968), LE MAITRE
(1970), ZLATIN; CHRISTOPHER (1973), ZLATIN; CHRISTOPHER (1974), TURLEY
(1981), LÓPEZ DE LACALLE et al. (2000)):
1. Resistance to plastic deformation needed to form chips due to the high strength of
titanium alloys that is maintained at elevated temperatures during machining;
2. High "coefficient of friction" at the tool-chip interface;
3. Generation of thin chips with reduced contact area with the cutting tool. This is
responsible for the generation of high stresses on the rake face of the tool during
machining;
53
4. Chip formation by adiabatic or catastrophic thermoplastic shear process. The
instability in the process of chip forming results in serrated (or segmented) chip;
5. Low modulus of elasticity of titanium alloys, which decreases quickly even at
moderate temperatures can lead to chatter, deflection and unwanted rubbing of the
tool to the freshly generated surfaces as well as bending of parts with thin walls and
with lack of precision in their finish;
6. High rate of work-hardening of titanium alloys;
7. High chemical reactivity of titanium at elevated temperatures (> 550ºC) with most
tool materials available and their consequent welding by adhesion onto the cutting
tool during machining leads to excessive chipping and/or premature tool failure and
poor surface finish;
8. The tendency of titanium alloys to ignite during machining at higher speed
conditions due to the high temperatures generated;
9. The variations in the cutting and in the axial components of the cutting strength,
cause chatter phenomena hence the need to have an stiff machine tool.
All these factors operating separately or in combination can cause rapid wear, chipping
or even catastrophic failure of the cutting tools.
2.13 Tool Materials for Machining Titanium Alloys
The machining system comprises of the cutting tool, the workpiece and the machine
tool. The cutting tool is the most critical in terms of machining productivity, therefore
selection of the right cutting tool material for a specific application is of crucial importance.
Other factors that contribute to the continuous development of cutting tool materials is the
development of new materials (generally with decreasing machinability) and the increasing
use of automated and numerically-controlled machine tools linked with systems which require
higher degree of reliability and predictability of cutting tools employed. Regarding to tool
cost, the cheapest tool material is not always considered the most economical nor is the most
expensive tool the best for a specific machining operation. The best tool material is one that
will maximise the efficiency and ensure accuracy at the lowest cost, in other words, the one
that will satisfy the requirements of a specific workpiece material (OKEKE, 1999).
54
It is known that cutting tool materials encounter severe thermal and mechanical shocks
when in machining titanium alloys. The combination of high temperature, high stresses,
strong chemical reactivity of titanium, formation process of catastrophic shear (irregular)
chips generated at and/or close to the cutting edge influence the wear rate and hence tool life.
Flank wear, nose wear, crater wear, notching, chipping and catastrophic failure are the typical
failure modes observed when machining titanium alloys. High Speed Steel (HSS) tool, coated
and uncoated cemented carbides, ceramic tools and ultra-hard materials such as
polycrystalline diamond and cubic boron nitride tools are the commercially available cutting
tool materials for machining of difficult-to-machine materials. High Speed Steel (HSS) tools
were widely used in earlier studies (DEARNLEY; GREARSON (1986), ZLATIN; FIELD
(1973), COLWELL; TRUCKENMILLER (1953), YANG (1970), CATT; MILWAIN (1968),
ZLATIN; CHRISTOPHER (1973)) on the machining processes of titanium at low cutting
speeds because of their limitations and lower cost relative to carbides. Although they are still
being used (KIM et al. (2001), SUN (1992)) for machining of titanium alloys at lower cutting
speeds conditions, their application have became reduced since cemented carbide tools have
proved their superiority in all machining processes of titanium alloys relative to competing
tool materials. Uncoated straight grade of cemented carbide (WC-Co) tools are the most
suitable for machining of titanium alloys due to their improved performance in terms of tool
wear and lower cost relative to other commercially available tool materials (ZLATIN; FIELD
(1973), JAWAID; CHE-HARON; ABDULLAH (1999), ZLATIN; CHRISTOPHER (1973),
ZLATIN; CHRISTOPHER (1974), KONIG (1979)). Inferior performance of coated carbides
tools (in terms of tool life) has been reported in many studies compared to uncoated carbide
grade tools when machining titanium alloys (CHRISTOPHER (1973), (KONIG (1979).
Recent developments in coating technology seem to demonstrate only negligible
improvement in machining titanium alloys. However, developments in advanced cutting tool
materials such as super-abrasive family, including PCD and CBN tools, have expanded the
application of these tools for the high speed machining of hard materials such as hardened
steels and titanium alloys despite their high cost (SHAW (1984), SECO TOOLS (2002b).
Some studies worldwide (SECO TOOLS (2002a), LEE (1981), HARTUNG; KRAMER
(1982), BHAUMIK; DIVAKAR; SINGH (1995), ZOYA; KRISHNAMURTHY (2000),
NABHANI (2001a), MACHADO et al. (2004), MAGALHÃES; FERREIRA (2004)) have
reported the superiority of PCD tools in terms of wear rate and hence longer tool life, when
machining of titanium alloys. Recent studies have reported that CBN tools can be used for
55
machining titanium alloys at higher cutting speeds despite the high reactivity of titanium
alloys with the tool materials (SECO TOOLS (2002a), HONG; MARKUS; JEONG (2001),
LEE (1981), HARTUNG; KRAMER (1982), BHAUMIK; DIVAKAR; SINGH (1995),
ZOYA; KRISHNAMURTHY (2000), NABHANI (2001a). Although ceramic tools have been
used for machining titanium alloys in several studies (YANG (1970), KONIG (1979), LEE
(1981), KOMANDURI (1989), LI; LOW (1994), KLOCKE; FRITSCH; GERSCHWILER
(2002), KOMANDURI; REED JR (1983)), they are not recommended for machining titanium
alloys because of their poor performance due to excessive wear rates as a result of the poor
thermal conductivity, relatively low fracture toughness and high reactivity with titanium
alloys (DEARNLEY; GREARSON, 1986).
2.13.1 Tool Materials Requirements
Tool selection process must be in accordance with the specification of the material to be
cut and cutting parameters. Since cutting tools used for machining titanium generally present
accelerated wear, due to extreme thermal and mechanical stresses close to the cutting edge, an
ideal cutting tool for machining titanium should possess the following requirements
(EZUGWU; WANG (1997), TRENT; WRIGHT (2000)):
i) hot hardness to maintain a sharp and consistent cutting edge at elevated temperature
and ability to withstand the high stress involved during machining;
ii) high resistance to abrasion wear in order to avoid alteration tool geometry caused
by the rubbing action;
iii) chemical inertness to minimise the tendency to react with titanium and chemical
stability to prevent the formation of a built-up edge;
iv) good thermal conductivity to minimise thermal gradients and thermal shocks;
v) high fracture toughness, which allows the insert to absorb forces and shock loads
during machining. If an insert is not sufficiently tough, then induced shock load
alone can lead to premature fracture of the cutting edge during machining;
vi) good fatigue resistance to withstand the chip segmentation process;
vii) high compressive, tensile and shear strength.
Szeszulski et al. (1990) consider wear resistance, chemical inertness and toughness the
main factors affecting the performance of cutting tools when machining aerospace alloys. The
reduction of hot hardness at elevated temperature conditions lead to the weakening of the
inter-particle bond strength and the consequent acceleration of tool wear. Table 2.5
56
(KRAMER, 1987) shows the softening temperature of commercially available cutting tool
materials while Figure 2.26 illustrates the influence of temperature on hardness of some
cutting tool materials used for machining aerospace alloys (ALMOND, 1981).
Table 2.5 - Softening points of tool materials (KRAMER, 1987).
Tool material Softening point temperature (
o
C)
High Speed steel 600
Cemented Carbide (WC) 1100
Aluminium Oxide (Al
2
O
3
) 1400
Cubic Boron Nitride (CBN) 1500
Diamond 1500
Figure 2.26 - Influence of temperature on hot hardness of some tool materials (ALMOND,
1981).
2.13.2 High Speed Steel (HSS) Tools
High Speed Steel (HSS) tool material was introduced by Taylor and White in 1906. The
chemical composition of initial HSS tool was: 0.67% C, 18.1% W, 5.47% Cr, 0.11% Mn,
57
0.29% V and Fe (balance), under proper heat treatment (TRENT; WRIGHT, 2000). The
advent of this allowed increase in metal removal rates and cutting speed up to 30 m min
-1
could be achieved when machining steel. This increase in cutting speed represented up to 6
times those achieved with carbon steels and low/medium alloy steels (most common cutting
tool materials employed in the 19
th
century). Subsequently a grade containing 18% W, 4% Cr
and 1% V was established and generally referred as the 18-4-1 or T grade. The room
temperature hardness of HSS tools is of the order of 850 HV (TRENT; WRIGHT, 2000).
There are several compositions of high speed steels in the market but three major grades can
be easily distinguished: the T-grade, M-grade and T/M grade (EDWARDS, 1993). The
tungsten grades of T series HSS tools contain 12-20% W with Cr, V and Co as principal
elements. These grades have a relatively wide hardness range and lower tendency to
descarburize, but they can readily acquire coarse-grained carbide particles resulting in lower
toughness. Molybdenum (Mo) based HSS tools (M series) were developed in the 1930´s and
are widely used because of the scarcity of tungsten. The Mo grades contains 3.5-10% Mo with
Cr, V, W and Co as alloying elements which tend to combine with carbon to form very
strongly bonded carbides with compositions such as Fe
3
(W,Mo)
3
C and V
4
C
3
(TRENT;
WRIGHT, 2000). The M-series HSS tools are tougher than the corresponding T-series HSS
tools. They are more widely used for drills and end mills. Increasing the Co content increases
the toughness but reduces hardness and therefore wear resistance of tool (BOOTHROYD;
KNIGHT (1989), EDWARDS (1993)). The third grade contains Co and can be either a T or
M series of HSS tools materials.
HSS tools can be employed in most of the common types of cutting tools including
single-point turning tools, milling cutters, taps, drills, reamers, thread cutting dies, gear
cutters, broaches, hobs and saws. Probably less the 10% of all turning applications are carried
out using HSS tools as the cutting tool material. The major area of application is drilling
which accounts for 80% of all drilling operations followed by milling (40% of the total
milling cutter market) (EDWARDS, 1993). These tools are also employed in the machining of
easy to cut materials and in operations where complex shaped tooling is involved. 350 HV is
often considered to be the maximum hardness of the workpiece which can be cut with HSS
tools. These materials generally are available in form of solid tools or in shape of inserts, both
could easily be reground when worn out. Various processes such as nitriding, carburising, ion
implantation and peening are used to optimise their properties.
58
High speed steel tools can be produced by three different processes: cast, wrought and
sintered (using powder metallurgy technique). Great progress was made in the 1970s with the
production of:
i) Coated HSS tools: layers of TiN, TiC, HfN and Al2O3 can be applied to the HSS
tools by either CVD (Chemical Vapour Deposition) or PVD (Physical Vapour
Deposition). High temperatures (up to 1000ºC) are required in the CVD process and
that may cause metallurgical changes to the parent HSS metal. The PVD process
operates at relatively lower temperature (500ºC to 600ºC) compared to CVD
process, making it more suitable for HSS tools (NEUMEYER; DULIS, 1976);
ii) Sintered HSS tools: powder metallurgy provide a microstructure consisting of a
uniform distribution of carbides throughout the matrix. As result, a substantial
increase in toughness can be achieved, ensuring for machining operations involving
impact and vibration. A HSS TiN coated indexable insert tooling was recently
designed for turning and parting operations. The substrate of the insert is powder
metallurgy HSS that ensured incorporation of sintered in chip breaker grooves
(NEUMEYER; DULIS, 1976).
The main advantages of using HSS tools are their high toughness relative to
commercially available cutting tools and reasonable cost and easy availability. Major
drawbacks of HSS tools include: hardness reduction at elevated temperature, limited wear
resistance and greater tendency for adhesion of the chip to the tool.
Generally purpose HSS tools grades suitable for machining of titanium alloys are the
M1, M2, M7 and M10 grades (THE METALS HANDBOOK, 1989). However, best results
are generally obtained with more highly alloyed grades including T5, T15, M33 and the M40
series. It was found that cutting speeds of 14 to 17 m min
-1
are reasonable for turning Ti-6Al-
4V alloy at a feed rate of 0.25 mm rev
-1
and at a depth of cut of 0.25 mm (THE METALS
HANDBOOK, 1989). The authors also reported that a M42 grade of HSS tool provided better
tool life than M2 or T15 grades. In other work a high energy electron beam was employed to
alloy the surface of M42 HSS tools with various boride powders including WB and MoB in
turning of both commercially pure titanium and Ti-6Al-4V alloy. It was reported that 35%
increase in cutting speed could be obtained with boride alloyed HSS tools compared with
unalloyed plain HSS tools (SUN, 1992). The authors also reported that below the cutting
speed limit the tool life of HSS tools increased by more than 7 times by electron beam
59
alloying with borides. This improvement was attributed to the lower flank and crater wear and
the higher hardness of the alloyed tools. Konig (1979) employed S18-1-2-10 HSS grade tools
in machining of Ti-6Al-4V alloy and reported that cutting speed achieved was much lower
than that used with cemented carbides (grade P) due to the rapid loss of hardness at elevated
temperatures above 600ºC. Similar performance was found when a M42 HSS grade and
cemented carbide tools were utilised in the experiments of Zlatin and Field (1973). Different
tool materials tend to have different responses to different wear mechanisms when machining
titanium alloys. The rapid loss of their hardness at elevated temperatures above 600ºC
subjects HSS tools to severe plastic deformation, which accelerate the rate of wear.
2.13.3 Cemented Carbide Tools
The introduction of cemented carbide tools in the early 1920s, in West Germany, by
Shroter led to a significant improvement in cutting speed capability and productivity due to an
unique combination of properties such as good wear resistance, strength and toughness.
Further development of cemented carbides began in the USA, Austria, Sweden and other
countries (TRENT; WRIGHT (2000), EDWARDS (1993), KALISH (1978), MARI;
GONSETH (1993)). Cemented carbides tools are produced by powder metallurgy technique
that involves the bonding of the tungsten carbide (56-93%WC) particles by a metal binder
cobalt (Co). Pure WC is comparatively brittle and Co is tough. The melting point of Co is
over 1400ºC, but there is WC-Co eutectic at about 1300ºC that facilitates liquid-phase
sintering. As a result of this, a strongly bonded, virtually 100% dense material is formed
KALISH (1978). A combination of these materials results in a compromise between wear
resistance and shock resistance according to the amount of Co used. Two factors affect the
properties of cemented carbide (WC-Co) (EDWARDS, 1993):
i) The Co content: increase in the Co content increases the toughness of a carbide but
reduces its hardness and therefore its wear resistance (BOOTHROYD; KNIGHT,
1989). Reduction of Co content reduces the toughness and increases the wear
resistance by increasing the hardness of a cemented carbide. Co content (in weight
percent) for cutting purposes generally ranges from 5% to 12%;
ii) The grain size of the WC: fine grain WC increases the hardness and therefore the
wear resistance for the wear resistance for a given Co content. Grain sizes of WC
ranges from around 0.5-5.0 µm (EDWARDS, 1993).
60
It has been reported (DEARNLEY; GREARSON (1986), LEE (1981), KOMANDURI;
REED JR (1983)) that hot compressive strength of cemented carbides depends on a
combination of the Co binder concentration and the WC particle size. The binding metal (Co)
easily dissolves in titanium thus carbides containing only small amounts of cobalt should be
used. Cemented carbides grades with low cobalt content and a fine grain size were found to
perform better than grades with higher cobalt content and/or a larger grain size when turning
titanium alloys (DEARNLEY; GREARSON, 1986). However, a lower Co content leads to
reduced rigidity of the cutting edge. These authors reported that the most favourable
compromise between a limited Co content and adequate rigidity of the cutting edge for
machining titanium is found in the carbides of the ISO grades K10 to K20. The grain size of
carbides tools affects their wear resistance. Mari and Gonseth (1993) reported that the fine
grain size (1.0 µm) of straight cemented carbides provided better resistance to deformation
than the coarse grain size (3.1 µm) below 800ºC (Figure 2.27).
0
400
800
1200
1600
2000
400 600 800 1000 1200
Temperature (ºC)
Flow stress, (MPa)
Fine grain size
Coarse grain size
Figure 2.27 - Flow stress measured at 0.6% strain during three point bending tests at a
constant strain rate in WC-11wt.%Co (MARI; GONSETH, 1993)
.
In contrast, the coarse grain size carbides showed a higher resistance to deformation at
higher temperature. Recent studies (JAWAID; CHE-HARON; ABDULLAH (1999), CHE-
HARON (2001)) were carried out to evaluate the performance of different grain sizes of
substrate of K10 straight uncoated carbide tools in machining of titanium, Ti-6242, alloy
61
under dry conditions at cutting speeds up to 100 m min
-1
and a constant feed rate of
0.25 mm/rev.
They reported that lower flank wear rate of the tools hence longer tool life is
due to the grain size of 1.0 µm compared to the finer grain size of 0.68 µm (Figure 2.28). The
authors attributed the greater wear rate of the finer grain size tools to increased solubility of
WC in titanium alloys as the surface area of tool particles exposed to solution wear increased.
0
2
4
6
8
10
12
60 75 100
Cutting speed (m/min)
Tool life (min)
883 grade
890 grade
Figure 2.28 - Tool live when turning Ti-6242 alloy with mixed uncoated carbide tools with
different grain sizes of substrates at a constant feed rate of 0.25 mm/rev: 0.68 µm (890 grade)
and 1.0 µm (883 grade) [adapted from JAWAID; CHE-HARON; ABDULLAH (1999)].
Carbide tools can be classified into six classes, according to ISO specifications which is
based on workpiece materials to be machined (INTERNATIONAL STANDARD, 2004):
1. P class: this is the class recommended for machining materials that produces long
continuous swarf: all grades of steel and cast steel except stainless steels with and
austenitic structure;
2. M class: this is recommended for machining stainless austenitic and
austenitic/ferritic steel and cast steel;
3. K class: this is mainly employed in machining of cast irons, such as grey cast iron,
cast iron with spheroidal graphite and malleable cast iron;
4. N class: this is recommended for machining aluminium and other non-ferrous
metals and non-metallic materials;
62
5. S class: this is the class recommended for machining difficult-to-machine materials,
including heat-resistant special alloys based on iron, nickel and cobalt, titanium and
titanium alloys;
6. H class: this class is recommended for machining hardened steel, hardened cast iron
materials and chilled cast iron.
Within each class of cemented carbides there are various categories (i.e. P01, P10, P20,
P30, P40, P50 and P05, P15 to P45), which are used for different machining operations.
Grades P01, M01, K01, N01, S01 and H01 are specific for precision finishing operations
because of increased wear resistance requirement under high speeds conditions. Grades P20,
M20, K20, N20, S20 and H20 are used for general purpose while P50, M40, K40, N30, S30
and H30 grades are recommended for roughing conditions due to increased toughness under
higher feed rates conditions.
Cemented carbides have a different designation in the USA, which will depend on the
type of work material to be machined. They can be categorised into two major divisions: C1
to C4 and C5 to C8. The former are suitable for machining cast iron, non-ferrous alloys and
non-metallic materials (non-steel materials) while grades C5 to C8 are suitable for machining
carbon steels and alloy steels.
Cemented carbides tools for machining applications are divided into two groups:
uncoated and coated carbides.
2.13.3.1 Uncoated Carbide Tools
The uncoated cemented carbides cutting tools are commonly used for all machining
processes of titanium alloys due to their improved performance in terms of tool wear and
lower cost relative to other tool materials. Two categories of uncoated carbide tools are
available for machining applications: straight and mixed grade carbides.
i) WC/Co carbides (straight grades): these grades are the original grades of cemented
carbides. The usual composition of the straight grade carbides is 94%WC and 6%
Co. In general, the best results in machining titanium alloys have been obtained
with the C-2 grades, represented by ISO K20 (DEARNLEY; GREARSON (1986),
ZLATIN; FIELD (1973), JAWAID; CHE-HARON; ABDULLAH (1999),
ZLATIN; CHRISTOPHER (1973), TRUCKS (1987)). However, satisfactory tool
lives (< 20 min) in turning titanium alloys were also obtained with ISO K10
63
uncoated straight carbides at cutting speeds below 60 m min
-1
(DEARNLEY;
GREARSON (1986), ZLATIN; FIELD (1973), (KOMANDURI; REED JR (1983).
Satisfactory tool lives were also obtained with ISO K10 uncoated straight carbide
tools when machining titanium alloys at cutting speeds of 60 m min
-1
and
122 m min
-1
with feed rate of 0.125 mm rev
-1
and depth of cut of 1.25 mm
(HARTUNG; KRAMER, 1982). All these authors reported that flank wear was
stable and did not contribute to tool failure until crater wear weakens the edge and
plastic deformation of the cutting edge accelerates wear at the tool flank. K10
uncoated straight carbides were also employed in turning of several grades of
titanium alloys at a relatively high cutting speed of 100 m min
-1
, feed rate of 0.1
mm rev
-1
and depth of cut of 0.5 mm (MOTONISHI et al., 1987). Recent studies
have employed K10 uncoated straight carbides in turning of titanium alloys at
relatively high cutting speeds ranging from 63 to 150 m min
-1
(EZUGWU;
OLAJIRE; WANG (2002), EZUGWU et al. (2004), EZUGWU et al. (2005),
KONIG (1979), MACHADO et al. (2004), MAGALHÃES; FERREIRA (2004),
KITAGAWA; KUBO, MAEKAWA (1997)). In most of these studies, the authors
observed that tool failure were due to adhesion and diffusion-dissolution wear
mechanism on the rake face, and attrition wear mechanisms on the flank face.
Plastic deformation on tool nose was responsible for tool rejection of uncoated
straight carbides when machining Ti-6Al-4V alloy at speeds in excess of 60 m min
-1
under conventional coolant supply (KONIG, 1979). Titanium reacts with most
cutting tool materials at elevated temperatures (> 550ºC) during machining
(MOTONISHI et al. (1987), KONIG (1979), HARTUNG; KRAMER (1982),
BROOKES; JAMES; NABHANI (1998)). Figure 2.29 shows crater wear rates of
various cutting tool materials when machining titanium alloys at a speed of
61 m min
-1
for 10 minutes. At such high temperature conditions titanium atoms
diffuse into the carbide tool material and react chemically with carbon present in the
tool to form an interlayer of titanium carbide (TiC) (KONIG (1979), HARTUNG;
KRAMER (1982)) which bonds strongly to both the tool and the chip to form a
saturated seizure zone, consequently, minimising diffusion wear mechanism as the
reaction is halted. The separation of the welded junction results in tool material
being carried away by the fast flowing chip. At lower speed conditions chemical
interactions between the carbide tool and titanium alloy is insignificant and wear is
64
caused mainly by mechanical and thermal fatigue as well as micro-fractures or
defects that may be present in the tool (EZUGWU; BONNEY; YAMANE, 2003).
Brookes, James and Nabhani (1998) utilised a quasi-static contact method in their
experiments to determine the temperature at which adhesion and welding developed
at the cutting zone that generated when machining titanium alloys with coated
carbide and CBN tools. They reported that the critical temperatures were 740ºC and
900ºC, for the carbide and CBN tool materials, respectively, and the nominal
contact pressures developed were approximately 0.23 GPa and 0.146 GPa in these
particular cases. It was also reported that fracture generally initiated in the bulk of
the harder tool material, rather than in the workpiece at the welded junction
interface. Figure 2.30 shows a SEM micrograph of exposed mixed carbide substrate
after fracture of the welded junction (NABHANI, 2001b). Uncoated straight
cemented carbide tools can also fail by abrasion wear mechanism as result of the
chipped hard particles flowing between the tool´s flank face and the newly
machined surface (BROOKES; JAMES; NABHANI, 1998). Figure 2.31 shows an
enlarged view of the flank face of a worn uncoated straight carbide tool that
experienced abrasion by carbide grains after turning of a titanium base, Ti-6242,
alloy under dry condition (JAWAID; CHE-HARON; ABDULLAH, 1999);
790
30
30
11
2.5
1.4
0
200
400
600
800
1000
Crater wear rate ( m/min)
Ceramic
(Al2O3)
CBN TiC-Coated
Carbide
TiN-Coated
Carbide
Uncoated
Carbide
PCD
Tool material
Figure 2.29 - Average crater wear rates of various tool materials in turning of Ti-6Al-4V alloy
at a cutting speed of 61 m min
-1
for 10 minutes [adapted from HARTUNG; KRAMER (1982]
65
.
Figure 2.30 - SEM micrograph of exposed mixed cemented carbide substrate after fracture of
the welded junction (NABHANI, 2001b).
Figure 2.31 - Flank face of a worn uncoated straight carbide tool showing abrasion by carbide
grains after turning titanium base, Ti-6242, alloy under dry condition (JAWAID; CHE-
HARON; ABDULLAH, 1999).
Mixed carbides (alloyed uncoated carbides grades): these grades have TiC, TaC or NbC
and other rare-earth elements such as ruthenium (Ru) added to the base composition of the
straight grade (WC-Co). Examples of typical compounds are WC/TiC/Co, WC/(Ta,Nb)C/Co,
WC/TiC/Ta(Nb)C/Co. Uncoated straight cemented carbides are the strongest and most wear
resistant but are subject to rapid cratering, especially when machining steels. TiC and TaC are
66
harder than WC and its inclusion improves the crater wear resistance of carbides
(BOOTHROYD; KNIGHT, 1989)) but have the disadvantage of lower strength. TiC usual
composition ranges from 5-25% by weight. The proportion of TiC added depends on the
cutting speed required. An increase in TiC composition generally reduces the toughness of the
carbide. Finishing operations generally are carried out at higher speeds for economic metal
removal. High speeds will cause cutting temperatures to increase and cratering will be more
pronounced. In this case high TiC content is required. When roughing operations are required
at lower cutting speeds conditions less TiC is required. TaC addition increases the hot
hardness of the tool, thus preventing plastic deformation of the cutting edge during machining
at high speed conditions. Mixed grade of carbides containing TiC (4% to 8%), Ta (Nb)C in
the range of 5% to 10% and Co (6% to 9%), with hardness range of 1450-1650 HV is also
recommended for machining superalloys. Due to the higher cost of TaC, it is often diluted
with up to 50% niobium carbide (NbC) without detracting from the performance of the
cemented carbide compound. WC/TiC/Ta(Nb)C/Co or Co-Ru cemented carbides can be
utilised for machining steels, where higher temperatures and high pressures are generated at
the cutting zone as result of a combination of high chemical instability and high temperature
properties. Ru can be substituted for the Co and increase in toughness, tool life and cutting
performance is achieved without alteration in hardness and other properties. Adhesion
diffusion-dissolution, attrition and plastic deformation wear mechanisms are responsible for
tool failure when machining with mixed uncoated cemented carbide tools. Figure 2.32 shows
the evidence of attrition wear and adhesion of the chips onto the nose of a worn mixed
((Ta,Nb)C) uncoated cemented carbide tool after machining Ti-6Al-4V alloy under
conventional coolant supply at a speed of 100 m min
-1
, a feed rate of 0.15 mm rev
-1
and a
depth of cut of 0.5 mm (EZUGWU et al., 2005).
67
Figure 2.32 - Evidence of adhesion of the chips on the nose of a mixed ((Ta,Nb)C) uncoated
(steel cutting grade) tool after machining Ti-6Al-4V alloy under conventional coolant supply
at a speed of 100 m min
-1
, a feed rate of 0.15 mm rev
-1
and a depth of cut of 0.5 mm
(EZUGWU et al., 2005).
2.13.3.2 Coated Carbide Tools
Coatings are used in cutting tools to provide improved lubrication at the tool-chip and
tool-workpiece interfaces, reduce friction and consequently lower temperatures at the cutting
edge. From a functional standpoint, chemical stability, hot hardness, and good adhesion to the
substrate are essential. Optimum coating thickness, fine microstructure, and compressive
residual stresses can further enhance coating performance (PRENGEL; PFOUTS;
SANTHANAN, 1998). Coatings must have a great resistance to wear in many aspects,
retaining their hardness when they are hot and reducing their affinity with the materials to be
machined. Additionally, coatings should possess low thermal conductivity to transfer less heat
to the substrate and low friction coefficient (LÓPEZ DE LACALLE et al., 2000). Other very
important characteristic that coatings must offer is a sharp and consistent cutting edge at
elevated temperature generated during machining, thus the coating should have the least
possible tendency to diffusion bond to the workpiece material (KLAPHAAK, 1987)
The most common group of coated tools consist of various combinations of titanium
nitride (TiN), titanium carbide (TiC), tantalum carbide (TaC), vanadium carbide (VC),
titanium carbonitride (TiCN), aluminium oxide (Al
2
O
3
), hafnium nitride (HfN), hafnium
carbide (HfC), zirconium nitride (ZrN) and chrome nitride (CrN) deposited in a single or
multilayer manner onto cemented carbides substrates. More recently new coatings with
68
improved properties have been introduced, including, titanium zirconium nitride (TiZrN),
titanium aluminium nitride (TiAlN), molybdenum carbide (Mo
2
C), titanium diboride (TiB
2
),
cubic boron nitride (CBN) and diamond-like coatings. These coatings are deposited on the
substrate either by the lower temperature (300-600ºC) Physical Vapour Deposition (PVD) or
higher temperature (600-1000ºC) Chemical Vapour Deposition (CVD) coating techniques.
Coatings deposited by CVD technique can vary in thickness from 5 to 20 µm, whereas PVD
coatings thickness are usually less than 5 µm. Diamond coatings can be deposited by CVD
techniques such as hot filament, microwave plasma and plasma torch methods at higher
temperature ranging from 700-1000ºC. Their coatings are generally in the range of 20-40 µm
(PRENGEL; PFOUTS; SANTHANAN, 1998).
When machining titanium alloys, tools wear out due to the high chemical reactivity with
titanium (HARTUNG; KRAMER (1982), BROOKES; JAMES; NABHANI (1998)). Thus,
the coatings should have high resistance to wear and high chemical stability to produce a
chemical barrier against heat, between the tool and the chip. Both TiC and TiAlN coatings
exhibit these properties. TiN is not a hard material, it provides a very low friction ratio in tool
cutting edges and greater resistance to crater wear (TURLEY, 1981). TiN is an even better
diffusion barrier than TiC but the latter has better abrasion resistance (EDWARDS, 1993).
TiC also improves hot hardness but, however, reduces fracture toughness. Layers of Al
2
O
3
,
TiN and TaC can be combined. The outer layers of TiN provide a low friction and chemically
inert surface for a new tool. As the TiN wears through, the additional wear resistant layer of
Al
2
O
3
is exposed (JAWAID, 1982). Al
2
O
3
coating can be either deposited by CVD or PVD
techniques. CVD Al
2
O
3
coating is most chemically stable of all hard coatings at all
temperatures, thus providing improved resistance to crater wear. CVD Al
2
O
3
coating is very
effective for the high-speed machining ferrous materials whereas PVD Al
2
O
3
coatings are soft
and unstable (PRENGEL; PFOUTS; SANTHANAN, 1998). The low thermal conductivity of
Al
2
O
3
coatings at high temperatures tends to concentrate the heat in the chips rather than the
tools. TaC addition increases the hot hardness of the tool, thus preventing plastic deformation
of the cutting edge at higher speed conditions. Hafnium nitride (HfN) coatings have high
hardness and chemical stability at elevated temperatures and a coefficient of thermal
expansion close to that of WC. These coatings have good resistance to abrasion wear,
cratering, and flank wear. TiB
2
is an emergent coating material exhibiting high hardness and
chemical inertness (PRENGEL; PFOUTS; SANTHANAN, 1998) deposited by CVD
technique. CBN coatings are chemically inert to hot iron, steel and oxidizing environments.
69
During oxidation, a thin layer of boron oxide is formed. This oxide provides chemical
stability to the coating, making them suitable for machining hard ferrous metals (50-65 HRC),
gray cast iron, high temperature alloys, and sintered powdered metals. CBN coatings also
have good adhesion to carbide substrate. Their thickness is generally restricted to 0.2-0.5 µm.
Diamond-like coatings have very high hardness, low friction coefficient, high thermal
conductivity and low thermal expansion coefficient. However, diamond has high affinity with
elements from group IVA to group VIIA of the periodic table what limits diamond coated
tools for machining non-ferrous alloys containing abrasive second-phase particles (i.e.
aluminium-silicon alloys) and for machining non-metallic materials that not react with carbon
(i.e. metal-matrix composites and fibre-reinforced plastics. Typical failure/wear modes for
diamond tools are combinations of oxidation, diffusion, micro-chipping and gross fractures
(PRENGEL; PFOUTS; SANTHANAN, 1998).
Coated carbides are currently employed in machining of titanium alloys at cutting
speeds similar to those for uncoated carbides, in the range of 56 to 150 m min
-1
(EZUGWU;
OLAJIRE; WANG (2002), VENUGOPAL et al. (2003), LÓPEZ DE LACALLE et al. (2000),
NABHANI (2001a), NABHANI (2001b), FITZSIMMONS; SARIN (2001)). Flank and nose
wears are the predominant failure modes when machining titanium alloys with coated
cemented carbide tools. Figure 2.33 shows typical flank wear plots for coated cemented
carbide tools when machining titanium alloy. When machining titanium alloys, multilayer
(TiC/TiCN/TiN) coatings of a mixed cemented carbide tool were rapidly removed from
cemented carbide substrate by adhesion wear mechanism taking place at the cutting edge
(Figure 2.34) (NABHANI, 2001b). The high adhesive forces are likely to result in the
plucking of hard particles from the tool. Thus erosion of the coating layer (s) exposes the
carbide substrate to extreme temperature at the cutting edge, consequently increasing crater
wear depth. A similar process of attrition and grooving wear can also develop on the flank
face, leading to deterioration in the machined surface. Ultimately, the combination of the
crater and flank wear undermine the integrity of the cutting edge leading to catastrophic
failure in extreme cases (NABHANI, 2001b). Adhesion wear mechanism also occurred in
tools coated with HfN when machining Ti-6Al-4V alloy (KONIG, 1979). Coatings of TiN,
TiC, Al
2
O
3
and HfN on both the rake and flank faces worn more rapidly than uncoated
straight grade carbides by either dissolution-diffusion or attrition wear mechanisms. Coatings
of TiB
2
and CBN tools were more resistant than others. In a recent study, Venugopal et al.
(2003) employed coated (TiB
2
) cemented carbide tools (P30 grade) in machining titanium
70
base, Ti-5Al-5Mo-2Sn-V, alloy under dry and cryogenic environments and reported that the
TiB
2
coating did not provide any benefit in all conditions tested. The TiB
2
coating was
removed by abrasion and attrition in both the rake and flank faces. They attributed this to
fluctuation of cutting forces during machining and the poor adhesion of the coating. Lee
(1981) carried out a study of performance of cooper-cooper iodide coating layers deposited to
a straight cemented carbide tools when machining titanium base, Ti-6Al-4V, alloy and
reported that increase in tool life was obtained when coated carbide tool was employed. The
author also reported that the coating was effective in improving seizure resistance, but the
ultimate performance of the coated tool was limited by the poor abrasion resistance of the
coating. Iodine is a lubricant additive for titanium in extreme pressure applications but has an
extremely corrosive characteristic, whereas cooper bonds well to cemented carbide tool
without affecting its surface. Cooper-iodide compound is one of more stable metal iodides.
Other study of evaluation of the performance of coated cemented carbide tools in turning Ti-
6Al-4V alloy was carried out by Walter, Skelly and Minnear (1993). In their studies cutting
edges of straight grade cemented carbides tools were implanted with halogen elements
(chlorine (Cl), fluorine (F), bromine (Br), iodine (I), sulphur (S)) and metallic elements
((indium) In, gallium (Ga) and tin (Sn)) by ion beam process to dose levels of 2x10
16
to 510
16
ions cm
-2
. These elements were expected to diffuse from the tool as a gas to react with the
workpiece material at high temperatures achieved during machining to provide a lubricant
film at the tool-workpiece. Improved tool life was obtained with both of halogen and metallic
implants. Chlorine and indium-implanted tools exhibited the best performance in terms of tool
life (tool life was increased by a factor of two). Additionally, the amount of titanium sticking
to the tool was reduced with coated-implanted tools compared to uncoated tools. The best
performance of chlorine implant was due to it reacting most readily with titanium, thus
providing the best lubrication to prevent or reduce sticking of titanium to the cutting edge of
the tool, leading to further failure of the tool. Since indium is insoluble in cobalt, so that at
least part of the indium would be present as liquid globules in the cobalt to supply liquid
metal to the interface during machining. This liquid metal would lubricate the interface by
preventing the sticking of titanium to the tool (WALTER; SKELLY; MINNEAR, 1993).
71
Figure 2.33 - Flank wear curves when machining Ti-6Al-4V alloy with coated (CrN and
TiCN) and a straight uncoated carbide tools (TURLEY, 1981).
(a)
(b)
Figure 2.34 - Worn surface of a multilayer (TiC/TiCN/TiN) coated mixed cemented carbide
tool showing remains of adherent metal layer (a) and enlarged view of the crater wear
showing smooth ridges with fine scoring in direction of chip flow (b) after machining
titanium base, Ti-5Al-4Mo-(2-2.5)Sn-(6-7)Si alloy (NABHANI, 2001b).
The coating deposition technique influences tool performance as the bonding strength of
the coating to the substrate varies with the processing technique employed. The flank wear
rate of multilayered CVD coated (TiCN + Al
2
O
3
) tools were found to be lower than that of the
single layered PVD coated (TiN) tools when milling Ti-6Al-4V alloy with tools between the
speed range of 55-100 m min
-1
and feed rate of 0.1-0.15 mm per tooth (JAWAID; SHARIF;
72
KOKSAL, 2000). The authors reported that these tools experienced excessive chipping at the
cutting edge and chipping and/or flaking on the rake face. The single PVD coated (TiN) layer
experienced higher flank wear rate. Additionally, it was observed that coating delamination,
adhesion of work material, attrition, diffusion, plastic deformation and thermal cracks were
the operating wear mechanisms of these tools (Figure 2.35). Delamination of the coatings can
be caused by either chemical reaction and/or crack propagation at the substrate interface,
which could be due to the difference in the thermal coefficient of expansion between the
coating matrix and the substrate (KONIG (1979), HARTUNG; KRAMER (1982)).
(a) (b)
Figure 2.35 - Coating delamination of PVD coated (TiN) carbide tool, grinding marks and
adhered material observed after 10 s (a) and adhesion of work material onto the flank face,
plastic deformation and cracks at the cutting edge after 20 s; (b) after face milling Ti-6Al-4V
alloy at cutting speeds of 100 m min
-1
and 50 m min
-1
and feed rates of 0.15 mm per tooth and
0.1 mm per tooth, respectively (JAWAID; SHARIF; KOKSAL, 2000).
2.13.4 Ultrahard (Superabrasive) Tool Materials
The term ultrahard, or alternatively superabrasive, tool materials refer to those tool
materials which possess a hardness of at least 3500 HV room temperature (ALMOND, 1981),
and encompasses diamond (both natural and synthetic) – also known as polycrystalline
diamond (PCD)) and cubic boron nitride (CBN). Throughout the 1990´s there has been
considerable growth in the use of PCD and CBN and polycrystalline cubic boron nitride
(PCBN) cutting tools. The automotive, aerospace and woodworking industries in particular
benefit from the higher levels of productivity, precision and consistency in manufacturing that
73
PCD and CBN/PCBN cutting tools can deliver (COOK; BOSSOM, 2000). These materials
currently have a well-established market and their properties and performance under severe
working conditions are generally considered to be outstanding and highly competitive.
2.13.4.1 Polycrystalline Diamond (PCD)Tools
Natural diamond is the hardest material known with hardness ranging from 6500-
12000 HV. The properties of diamonds which are suitable for cutting tool materials include
low coefficient of friction, high thermal conductivity, non-adherence to most materials, ability
to form a sharp cleavage edge and high abrasion wear resistance ensuring that their cutting
edge is maintained throughout most of their useful life. Low compressibility and thermal
expansion provide dimensional stability, assuring the maintenance of close tolerances and the
generation of high surface finish. However, due to its extreme brittleness, natural diamond is
prone to premature failure as well as being very difficult to shape into cutting tools.
Machining with diamond cutting tools is generally carried out at high speed and low feed rate
conditions to prevent catastrophic failure. Another drawback of sintered diamond is its high
chemical reactivity with ferrous materials at elevated temperatures and to revert, at high
temperature (about 700ºC), to graphite and/or to oxidise in air (DE GARMO; BLACK;
KOHSER, 1999).
Diamond was successfully synthesised in 1958 (EDWARDS, 1993) by subjecting
carbon (graphite) to very high temperatures (1500ºC) and ultrahigh pressures (6000 MPa) in
the presence of a catalyst or solvent metal (manganese, iron, cobalt, nickel, palladium,
platinum and their alloys). This produces fine diamond particles (grain size 1-30 µm), which
are then fused together at the appropriate sintering conditions of pressure and temperature to
produce polycrystalline diamond. PCD became available as cutting tool material in the mid
1970s. The manufacturing route is basically the same as the used to produce diamond grit,
however the temperature and pressure involved are not as high.
Polycrystalline diamond (PCD) tools consist of a thin layer (about 0.5-1.5 mm) of fine
grained diamond particles sintered together and metallurgically bonded to a tungsten carbide
(WC) substrate which provides adequate toughness and strength. The sintered diamond tools
are then finished to shape, size and accuracy by laser cutting and grinding. Typical properties
of PCD are: hardness from 7000 to 10000HV, fracture toughness of approximately 8 MPa
m
0.5
and thermal conductivity around 600 Wm
-1
K
-1
. The main limitation of PCD tools is that
they are not generally suitable for machining ferrous metals due to low chemical stability at
74
elevated temperatures. Diamond reverts to graphite if exposed to temperatures greater than
750ºC for periods longer than 1 minute. Problems of degradation are such that it is not
recommended for the machining of ferrous alloys and special care must be taken during
brazing. PCD tools are mainly used for machining abrasive non-metallic materials (such as
carbon, ceramics, graphite, fibreglass and its composites, reinforced plastic and rubber), non-
ferrous metals (such as aluminium alloys, brass, bronze, cooper, zinc and their alloys) and
tungsten carbides. PCD tools are widely used for milling, turning, boring, threading and other
operations in the mass production of many aluminium alloys, including free machining
aluminium alloys and high silicon aluminium alloys because the very long tool life without
regrinding can reduce costs (TRENT; WRIGHT, 2000).
PCD grades are usually differentiated by their diamond particle size. Increasing grain
size increases abrasion resistance and decreasing grain size increases edge quality (under
comparable conditions). Coarse grain size PCD grade is recommended for machining under
harsh conditions. Medium and fine grain size PCD grades are recommended for machining
under moderately harsh conditions because their resistance is a less important factor and
factors such as edge quality/surface finish must be considered. However, Cook and Bossom
(2000) carried out a study of evaluation of performance of various PCD grades in milling
ceramic impregnated layer and reported that relationship between the PCD grain size and its
abrasion wear resistance is not linear (Figure 2.36) and that tool life dropped if the grain size
is in excess of 25 µm. When using an ultra coarse grain size of 75 µm, the tool life obtained
was considerably lower than those obtained with coarse and medium grain sizes. Similar
results were obtained by BAI et al. (2004) in their study evaluating the effects of diamond
grain size (2-75 µm) on flank wear of PCD tools when machining laminated flooring with
Al
2
O
3
overlay. The authors reported that smallest flank wear occurred on the tool with 25 µm
grain size. In both studies, the worst performance of ultra coarse grain size was attributed to
increased brittleness and rougher cutting edge of coarse grain size, which make them break or
abrade, adversely affecting overall performance.
75
Figure 2.36 - The performance of various grades of PCD tools when milling ceramic
impregnated surface of a flooring board (HPL) (COOK; BOSSOM, 2000).
There have been some reports that PCD tools were successfully used for machining
titanium base alloys (SECO TOOLS (2002a), LEE (1981), HARTUNG; KRAMER (1982),
BHAUMIK; DIVAKAR; SINGH (1995), ZOYA; KRISHNAMURTHY (2000), NABHANI
(2001a), MACHADO et al. (2004), MAGALHÃES; FERREIRA (2004), (NABHANI,
2001b)) despite their higher cost, over 8900% higher than cemented carbides (SECO TOOLS,
2002b). A recent study (MACHADO et al., 2004) on evaluation of various cutting tools in
finish turning of titanium, Ti-6Al-4V, alloy at high cutting speeds of 110 to 500 m min
-1
reported that PCD tools gave the longest tool life compared to cemented carbides, CBN and
ceramic tool materials. A speed of 175 m min
-1
was reported as the optimum speed to
machine titanium under conventional coolant flow. Superior performance of PCD tools in
terms of tool life and surface finish generated relative to CBN and carbide tools was also
reported by Nabhani (2001b) when machining titanium base, TA48, alloy at a cutting speed of
75 m min
-1
, a feed rate of 0.25 mm rev
-1
and a depth of cut of 1.0 mm under dry condition.
The longer tool life recorded with PCD tools was attributed to interdiffusion of titanium and
carbon taking place at the interface tool-workpiece, which resulted in the formation of a
titanium carbide layer on the rake face of the tool. There was a workpiece layer adhering to
the substrate of PCD tool which prevented further diffusion of the tool material into the chip
(Figure 2.37). Nabhani (2001b) also reported that the critical interfacial temperature for
adhesion to occur when machining titanium alloy with PCD is 760ºC and the nominal
pressure is about 0.142 GPa whereas for carbides the temperature is 740ºC and pressure of
76
0.23 GPa. Dearnley and Grearson (1986) reported that the solubility limit of the tool material
in the workpiece is the main factor involved in the dissolution-diffusion process, which
determines the magnitude of the concentration gradient in the shear zone, and hence the
diffusion flux. This phenomenon was also confirmed by Hartung and Kramer (1982),
suggesting that the wear rate of tool materials which maintains a stable reaction layer is
limited by the diffusion rate of tool constituents from the tool-chip interface. At equilibrium,
this flux will be equal to the diffusion flux of tool constituents through the reaction layer.
These authors found that PCD and uncoated cemented carbide tools showed evidence of the
formation of stable reaction layer, and were the most wear resistant materials compared to
coated cemented carbide and CBN tool materials when machining Ti-6Al-4V alloy.
Figure 2.37 - Formation of strongly adherent layer on the rake face of a PCD tool after
machining titanium base, Ti-5Al-4Mo-2Sn-6Si alloy under dry condition (NABHANI,
2001b).
2.13.4.2 Cubic Boron Nitride (CBN) Tools
Cubic boron nitride (CBN) is the hardest material next to diamond (4000 - 6000 HV).
CBN is not a naturally occurring compound. The synthesis of CBN from hexagonal boron
nitride (HBN) was first announced in 1957 following developments on the synthesis of
diamond. CBN powder is manufactured by subjecting HBN to extremely high pressure
(6000 MPa) and high temperature (around 1400ºC). CBN can be monocrystalline or
polycrystalline and is used for cutting metal when cemented carbide becomes limited in the
cutting speeds that can be employed. Table 2.6 shows the mechanical and physical properties
and capital cost of a typical high concentration CBN/PCBN tool compared with other cutting
77
tool materials mentioned previously (ABRÃO, 1995). It can be seen that CBN/PCBN tools
retain hardness at high temperatures of 1000ºC with very good fracture toughness, properties
that combine with its relatively low solubility in iron to enable CBN/PCBN to machine hard
ferrous alloys and some cobalt and nickel superalloys faster than other tool materials. CBN is
also employed in machining ferrous materials of hardness of 450 HV or above (tool steels,
case hardened steel, chilled cast iron) and high temperature alloys (nickel or cobalt base) of
340 HV and above (RICHARDS; ASPINWALL (1989), WATERS (2000)). The cost of a
CBN is about 6500% higher that of a carbide tool (SECO TOOLS, 2002b). Its use has been
restricted to finish machining operations in order to effectively compete with grinding which
is an expensive process for generating complex surface. Commercial CBN tool blanks/inserts
are available in two major grades, namely: CBN
-H
(High CBN content) and CBN
-L
(Low CBN
content). The CBN
-H
grade is generally harder and has higher fracture toughness. The CBN
-L
grade contains in addition to CBN either titanium carbide (TiC) or titanium nitride (TiN) to
offer high wear resistance, with reduced fracture toughness. It is generally recognized that low
CBN content tools have longer life than those of high CBN content tools in machining of
difficult-to-machine materials, especially titanium alloys, due to the ability of the lower CBN
content grades to retain their cutting edge for longer period (RICHARDS; ASPINWALL
(1989), WATERS (2000), HUANG; LIANG (2003)). In addition to CBN content, the
performance of CBN tools depends on the microstructure of binder phase, manufacturing
process employed and method of grinding the tool geometry. CBN tools also differ by types
of structure: the solid type (indexable insert type) and the layered type (brazeable blank type).
The former consists of polycrystalline compact only, whereas the later comprises a layer of
polycrystalline CBN in WC substrate.
Notching, chipping and premature tool failure are the dominant failures modes of CBN
tools due to their brittle nature. Premature failure of CBN tools can be attributed to the
fracturing of the unsupported cutting edge caused by attrition wear (BHAUMIK; DIVAKAR;
SINGH, 1995). Recent studies have reported that CBN tools can be used for machining high
temperature superalloys, especially titanium-alloys, at a higher cutting speeds
(up to 350 m min
-1
) despite the high reactivity of titanium-alloys with the tool materials in
addition to their relatively high cost (SECO TOOLS (2002a), HONG; MARKUS; JEONG
(2001), HARTUNG; KRAMER (1982), BHAUMIK; DIVAKAR; SINGH (1995), ZOYA;
KRISHNAMURTHY (2000), NABHANI (2001a)). Owing to their high hardness and high
melting point, CBN tools can withstand the heat and pressure developed during cutting
78
without compromising surface integrity of the machined component because of their ability to
maintain a sharp cutting edge for longer period (SECO TOOLS, 2002a). Kramer (1987)
reported that CBN tools retain their strength at temperatures in excess of 1100ºC while
cemented carbide tools encounter plastic deformation at this temperature range. This
temperature is also that in which titanium can react with nitrogen in the CBN when
interatomic diffusion of titanium and CBN accelerate significantly. This process was
confirmed by Nabhani (2001a) when machining titanium-base, Ti-5Al-4Mo-(2-2.5)Sn-(6-
7)Si, alloy with CBN and PCD tools where the cutting edge of the tool was bonded to the
underside of the chip as shown in Figures 2.38 (a) and (b). CBN tools are recommended for
finish machining of titanium alloys at a speed of about 150 m min
-1
and a depth of cut in
excess of 0.5 mm (SECO TOOLS, 2002a).
(a) (b)
Figure 2.38 - (a) Section through ‘quick-stop’ specimen showing part of CBN tool adhering to
underside of chip (100x), (b) close-up view of Fig. 2.38(a) (200x) (NABHANI, 2001a).
Cutting speed plays an important role in the performance of CBN tools. Increase in
cutting speed leads to increase in cutting temperature as well as increasing the intensity of
chemical interaction between the tool and the workpiece materials, which subsequently
affects tool life. As the cutting temperature increases, seizure of the chip occurs everywhere
on the tool face. This forms an adherent layer which becomes saturated with tool particles and
serves as a diffusion boundary layer, thus reducing the rate of transport of tool materials into
the chip and consequently the wear rate. Since the diffusivity rate increase exponentially with
temperature, further increase in cutting speed beyond the speed for minimum wear produces
rapid increase in the wear rate. In a study carried out by Dearnley and Grearson (1986) on the
79
machinability of titanium base, Ti-6Al-4V, alloy with various cutting tool materials, it was
reported that CBN tools showed high wear rates with rake and flank surfaces smoothly worn
at cutting speeds in excess of 100 m min
-1
, hence they were not considered suitable for
practical application. At cutting speeds lower than 100 m min
-1
crater wear was irregular,
suggesting attrition, and the wear rates exceeded those observed with uncoated straight grade
carbides. In other study on evaluation of performance of CBN tools in machining of titanium
alloy, Zoya and Krishnamurthy (2000) reported that the occurrence of the increased
temperature depends on the cutting speed employed. They also reported that the increase in
wear rate occurred around 700ºC, indicating the critical temperature constraining the
performance of CBN tools when machining titanium alloys. When machining at a relatively
low speed of 150 m min
-1
, a feed rate of 0.05 mm rev
-1
and a depth of cut of 0.5 mm, tool
wear was associated with localized chipping of the cutting edge (Figure 2.39 (a)), possibly
due to tool-tip oscillations. With increase in cutting speed and consequently increase in
cutting temperature, cutting tool experienced crater wear and severe nose wear (Figure
2.39(b)). Outstanding performance of wurtzite boron nitride (wBN)-CBN composite cutting
tools relative to PCD and cemented carbide tools was observed when machining titanium-
base, Ti-6Al-4V, alloy at a cutting speed of 75 m min
-1
, a feed rate of 0.1 mm rev
-1
and a
depth of cut of 0.5 mm (BHAUMIK; DIVAKAR; SINGH, 1995). This was attributed to a
stable interface layer formed on the composite tools, which protected the rake face as well as
helped in easy shearing of the work material. This layer also inhibits further diffusion and
dissolution, thereby preventing the chip from adhering to the tool and consequently reducing
the rate of wear. Brookes; James and Nabhani (1998) found that CBN tools showed lower
wear rate and provided better surface finish than coated carbide tools when machining
titanium, Ti-5Al-4Mo-(2-2.5)Sn-(6-7)Si, alloy at high cutting speeds. They attributed this in
part to the relatively low solubility of boron and also to their higher hardness and melting
temperature. Satisfactory tool life and surface finish have been reported when dry machining
titanium-base, Ti-5Al-4Mo-(2-2.5)Sn-(6-7)Si, alloy with ultra-hard materials (PCD, CBN)
and coated carbide tools at a cutting speed of 75 m min
-1
, a feed rate of 0.25 mm rev
-1
and a
depth of cut of 1.0 mm (NABHANI, 2001a). Average flank wear was the dominant failure
mode. These results show that ultra-hard tool materials can be used in dry machining of
titanium alloys at lower speeds and at relatively lower feed rates. Despite encouraging
performance of CBN tools, their application in dry machining of titanium
alloys is still
questionable due to their higher cost compared to carbides and ceramic tools.
80
(a)
(b)
Figure 2.39 - (a) A typical scanning electron micrographs of worn-out edges: (a) cutting
temperature of 734ºC, (b) cutting temperature of 900ºC (ZOYA; KRISHNAMURTHY,
2000).
2.13.5 Ceramic Tools
Although ceramic tool materials were in use thousands of years ago it was not before
the early 20
th
century that they became important again. During and after Second World War,
the scarcity of tungsten, and later the development of stronger alloys, compelled material
scientists to improve the mechanical properties of ceramic cutting tools by varying the
manufacturing routes and trying different additives. The brittleness of ceramic tool materials,
coupled with insufficient rigidity and lack of power within the machine tool, hindered their
application and usage during their early development. However, these problems have been
overcome by developing new advanced ceramics and more powerful, rigid and stable machine
tools. Ceramic tools are hard and can retain their hardness at elevated temperature conditions
generated when machining at high speed conditions. They also exhibit better chemical
inertness and oxidation resistance than cemented carbides. However, the inadequate fracture
toughness of ceramic tools relative to HSS tools and cemented carbides is the major
disadvantage in their use. This makes them susceptible to mechanical and thermal shock
during machining (NORTH, 1986). Ceramics have high melting point and the absence of a
secondary binder phase, like carbides, prevents them from softening under high speeds
conditions (CHAKORABORTY; RAY; BHADURI, 2000). Ceramic cutting tools are mainly
classified into two groups:
a) Alumina-base (Al
2
O
3
) ceramics (pure oxide, mixed oxides and silicon carbide (SiC)
whisker reinforced alumina ceramics);
81
b) Silicon nitride-base (Si
3
N
4
) ceramics.
2.13.5
The pure oxide (Al
2
O
3
) are composed of fine Al
2
O
3
grains which are sintered
87). They possess high hardness, good wear resistance and
excel
he mixed oxide ceramics are obtained by the addition of high percentages of TiB
2
or
re Al
2
O
3
ceramic. Depending on the manufacturer, mixtures
of Al
.1 Pure Oxide Ceramics
(EZUGWU; WALLBANK, 19
lent chemical stability but they lack toughness. The toughness of alumina-base ceramics
has been improved by the addition of zirconium oxide (ZrO
2
) because of its phase
transformation characteristics without affecting their wear resistance. Typical applications of
pure oxide ceramics are both finishing and rough cutting of cast materials. However, pure
oxide ceramic tools have not been effective in machining aerospace alloys due to their poor
thermal shock resistance and low fracture toughness at elevated temperatures.
2.13.5.2 Mixed Oxide Ceramics
T
TiC and/or TiN particles to the pu
2
O
3
with selected additives may be either compacted isostatically, followed by sintering
at high temperature, or the powders sintered at high temperature under pressure (LI; LOW,
1994). Magnesium oxide is often added to pure Al
2
O
3
ceramic in small amounts to inhibit
grain growth during sintering. Additional additives may consist of small amounts of other
materials such as chromium oxide, titanium oxide, nickel oxide or refractory metal carbides
(WHITNEY, 1983). The purpose of these additives is to achieve certain properties in the tool,
particularly increased mechanical strength. Mixed oxide ceramics generally possess improved
fracture toughness, hardness and thermal shock resistance (EDWARDS, 1993). The addition
of TiC to alumina makes the material more difficult to sinter and so hot pressing is generally
employed (RICHARDS; ASPINWALL, 1989); TiC improves hot hardness but, however,
reduces fracture toughness. The use of TiN enables the material to be cold pressed and
sintered. Introduction of TiC makes the material to be black in colour whilst the Al
2
O
3
+ TiN
is dark brown. These modified alumina ceramics are variously known as “mixed”, “black” or
“hot pressed” ceramics. The mixed ceramics are not only thermally tougher but have better
retention of the hardness level at elevated temperatures relative to the pure oxide ceramics.
The main disadvantage in their use is that the mechanical toughness is inferior to that of the
pure oxide ceramics (RICHARDS; ASPINWALL, 1989).
82
2.13.5.3 Whisker Reinforced Alumina Ceramics
Whisker reinforced (Al
2
O
3
+ SiC
w
) ceramic tools consist of the basic alumina matrix
streng ened mechanically with at least 25 wt.% fibres of silicon carbide (SiC
w
) (LI; LOW,
nforcement is to improve the strength and
tou
ceramics tools has long been recognised as the toughest
ceram materials. Improvement in hardness, abrasive wear resistance and chemical inertness
utting tools. These cutting tools are manufactured by
sinter
th
1994). The main advantage of the whisker rei
ghness of the compact through maximising its crack deflection and whisker pull-out
mechanisms at elevated temperature condition. This tool grade also possesses low thermal
expansion coefficient and high thermal conductivity. The combination of these two factors
improves thermal shock resistance of the whisker reinforced alumina ceramics compared to
non-reinforced alumina ceramics.
2.13.5.4 Silicon Nitride-base Ceramics
Silicon nitride-base (Si
3
N
4
)
ic
have produced a material suitable for c
ing Si
3
N
4
crystals with an inter-granular glass phase of SiO
2
in the presence of Al
2
O
3
,
yttria (Y
2
O
3
) and magnesium oxide (MgO). There are two forms of Si
3
N
4
base ceramic tools
for machining purposes: α- Si
3
N
4
and β’- Si
3
N
4
. The α- Si
3
N
4
is produced by nitriding of
silicon at temperatures up to 1300ºC with small amounts of Y
2
O
3
and higher aluminium
content (CSELLE; BARIMANI, 1995). The β’- Si
3
N
4
is a covalent solid which contains
negligible amount of oxygen. It was initially produced by hot pressing at pressure in the range
of 7.6 to 17.8 MPa and a temperature in the range of 1650 to 1775ºC, until a density of at least
3.25x10
3
kg m
-3
is obtained (LI; LOW, 1994). Their properties can differ for each individual
grade depending on its composition, processing route and to a larger extent on the type and
quantity of the intergranular phase (RICHARDS; ASPINWALL, 1989). One of most
important Si
3
N
4
base ceramic is the silicon aluminium oxynitride (also referred to SIALON).
SIALON ceramic is the result of substitution of oxygen (O
2-
) by nitrogen (N
3+
) in the β’-
Si
3
N
4
crystal provided that aluminium (Al
3+
) is simultaneously substituted for silicon (Si
4+
) to
maintain a charge neutrality. This material has the same crystal structure as β’- Si
3
N
4
, and
similar physical properties, but better chemical properties because of the chemical substitution
(LI; LOW, 1994). SIALON possesses excellent thermal shock resistance due to their low
thermal expansion coefficient as well as high thermal conductivity, so that the stresses
between hot and cool parts of a cutting tool insert are minimised, giving good thermal shock
resistance (MOMPER (1987), RICHARDS; ASPINWALL (1989)). The relatively high
83
fracture toughness of Si
3
N
4
base ceramic tools is attributed to a fine microstructure containing
elongated grains and the fact that cracks are controlled by the pulling out of grains form the
matrix. This property make them suitable for rough machining of nickel-base alloys at higher
speed conditions as well as rough turning of grey cast iron with strong interruptions of cut,
rough milling of grey cast iron and turning of materials having a high nickel content
(MOMPER (1987), RICHARDS; ASPINWALL (1989)).
Although ceramic tools have been used for machining of titanium alloys in several
studies (DEARNLEY; GREARSON (1986), LEE (1981), KOMANDURI (1989), LI; LOW
(1994), KLOCKE; FRITSCH; GERSCHWILER (2002),
KOMANDURI; REED JR (1983))
they a
re not commercially recommended for machining titanium alloys because of their poor
performance due to excessive wear rates as a result of the poor thermal conductivity,
relatively low fracture toughness and high reactivity with titanium alloys (Figure 2.29)
(HARTUNG; KRAMER, 1982). Lower tool lives (< 30 seconds) were recorded when
machining titanium-base, Ti-6Al-4V, alloy with various grades of ceramics (SIALON, Al
2
O
3
+ TiC and Al
2
O
3
+30ZrO
2
) at a cutting speed of 75 m min
-1
compared with cemented carbide
tools (DEARNLEY; GREARSON, 1986). It was reported from this study that all the ceramics
exhibited severe notching at depth of cut and the cutting edges of most ceramics revealed
smoothly worn rake faces as a result of a dissolution-diffusion and attrition wear mechanisms.
SIALON was the most resistant to rake face wear. Smooth and irregular/uneven worn
surfaces were observed, but it was, however, prone to severe notching and crater wear, thus
not reliable in machining of titanium alloys. Poor performance of SIALON tool was also
confirmed by Komanduri and Reed (KOMANDURI; REED JR, 1983) during a study of
evaluation of various grades and geometries of cemented carbides and SIALON tools for
machining titanium alloys at cutting speeds up to 183 m min
-1
(Figure 2.40). They reported
that the SIALON tools failed almost instantaneously (in less than 15 seconds) with high
localised flank wear and cratering, which was followed by notching and catastrophic failure,
thus not suitable for machining titanium alloy. Lee (1981) reported that less than one minute
tool life was recorded when a pure oxide alumina-base and zirconium-base (Al
2
O
3
+ZrO
2
)
alumina ceramics were utilised for machining titanium alloy at a cutting speeds in excess of
122 m min
-1
. The author attributed the poor performance of ceramic tools to their poor
thermal shock resistance and low strength. Other study employed pure oxide alumina-base
and SiC whisker reinforced ceramic cutting tools in machining of Ti-6Al-4V alloy at cutting
speeds up to 600 min
-1
, feed rate of 0.1 mm rev
-1
and depth of cut of 1.0 mm (KLOCKE;
84
FRITSCH; GERSCHWILER, 2002). The longest tool life of 4 minutes was recorded when a
cutting speed of 15 m min
–1
was employed, suggesting that these tools are not suitable for
machining titanium alloys. All the tools were subjected to notch wear, flank wear and crater
wear.
Figure 2.40 - Variation in uniform flank wear with cutting time for the turning of Ti-6Al-4V
(hardness, 36 HRC), showing reduced tool wear with the new geometry (cutting speed,
122 m min
-1
; feed rate, 0.23 mm rev
-1
unless otherwise indicated; depth of cut, 1.52 mm; tool
ir physical, mechanical
and chemical properties (NORTH, 1986). Although conventional micron-grain size ceramic
g performance when machining titanium alloys, advances in
nano-
and morphology of the grain and pore as well as the distribution of impurities in the tool
SNG432 (SCEA, 15º): curve A, SIALON (Kyon 2000) with clearance angles of 17º (localised
wear, 0.889 mm; edge fracture; crater) and 5º (localized wear, 1.321 mm; fracture; crater);
curve B, SIALON (Kyon 2000) with a clearance angle of 5º and a feed rate of
0.127 mm rev
-1
; curves C and D, cemented carbide (Carboloy grade 999) with clearance
angles of 5º and 17º, respectively (KOMANDURI; REED JR, 1983).
2.13.5.5 Nano-grain Ceramic
Performance of ceramic tools during machining depends on the
tools did not show encouragin
technology have generally led to improved mechanical properties of cutting tools,
especially ceramics.
Mechanical properties, such as the hardness, strength and density are related to the size
85
substrate. The use of finer grained substrate generally produced densely packed compact tools
after sintering. The fracture modes of ceramics vary with the grain size, which in turn affect
the fracture toughness. Increasing the pore size and inter-granular porosity tends to reduce
fracture toughness. Wear resistant capabilities of ceramic cutting tool materials are also
associated with their grain size. Nano-grain ceramic cutting tool materials have been
developed to further enhance their machining performance. The fine grained ceramics also
called nano-ceramic possess high specific area, improved thermal shock resistance, improved
hardness and superplastic behaviour as well as low thermal expansion coefficient (4.5x10
-6
ºC
-1
) (VAβEN; STÖVER, 1999). Powdered nano-ceramics is defined as a material where the
major phase (or at least one constituent) has a grain size in the nanometer range (Figure 2.41)
(VAβEN; STÖVER, 1999).
Figure 2.41 - TEM micrograph of HIPed (Hot Isostatic Pressing) nanophase SiC sample with
a density of 97% TD (Theoretical density) (VAβEN; STÖVER, 1999).
Powdered nano-ceramics have good formability properties due to their superplasticity
elongation in tensile
deformation (XIE; MITOMO; ZHAN, 2000). The properties of nano-ceramics to a greater
exten
behaviour and hence can be formed in complex geometry. Superplasticity is defined by the
ability of polycrystalline solids to exhibit exceptionally high levels of
t depend on the processing technique employed. Despite the fact that nano-ceramics and
conventional ceramics generally have the same composition, the production process of the
former require high temperature and pressure for efficient sintering in order to reduce the
number of agglomerates in their powders, thus ensuring a denser phase powder than
conventional ceramics (ZHANG; ZHU; HU, 1996). The sintering pressures can exceed 8 GPa
and temperature in excess of 1000ºC. The density of nano-ceramic powder has significant
86
influence on the mechanical properties such as ductility at low temperature, superplasticity at
elevated temperatures and hardness. It is therefore anticipated that this technique will produce
cutting tools with improved wear resistance, hardness and toughness relative to conventional
ceramics (VAβEN; STÖVER, 1999). Kear et al. (2001) compared the abrasive wear
resistance between micron and nano-grained ceramic particles of TiOB
2
B and concluded that the
latter presented about 10% more wear resistance than the former. They also observed a
reduction of about 50% in the friction coefficient, surface roughness (R
B
a
B) between 20-50 nm,
the development of surface plasticity and improved toughness for the nano-grained TiO
B
2
B.
Bhaduri and Bhaduri (1997) also investigated the toughness behaviour of Al
B
2
BOB
3
B-ZrOB
2
Bnano-
ceramic and observed a slight drop in hardness with improved toughness in relation to
conventional ceramics. The improved hardness observed with reduction in grain size of SiC
nano-ceramics (VAβEN; STÖVER, 1999) was attributed to lesser of porosities within the
phases with higher density of particles. Furthermore, they observed an excellent thermal
shock resistance in materials such as SiC/C-composites with a SiC matrix grain size lower
than 200 nm, in spite of the low thermal conductivity relative to conventional ceramics. Kim
(1994) verified that increasing the pore size and inter-granular porosity lead to reduction of
fracture toughness in ceramic tools. Ezugwu; Bonney and Olajire (2002) observed that
micron grain whisker reinforced alumina ceramic tool out performed the nano-grain ceramic
tools grades (Al
B
2
BOB
3
B and SiB
3
BNB
4
B) at cutting speed range of 230-270 m minP
-1
, feed rate range of
0.125-0.15 mm rev
-1
under conventional coolant flow when machining Inconel 718. Silicon
nitride (Si
B
3
BNB
4
B) base nano-ceramic also gave the worst performance in terms of tool life due to
high nose wear rates attributed to the softening of tool materials and consequent weakening of
their bond strength when machining at higher speed conditions. A very interesting fact is that
nano-grain ceramic tools are more stable in terms of recorded tool life when machining at
higher cutting speeds in excess of 230 m min
P
-1
P
and at a feed rate of 0.125 mm rev
-1
.
2.14 Cutting Fluids
There are many ways to ensure effective and efficient metal cutting. The use of
a cutting
fluid is one the options and when properly chosen and applied, can be successful. Cutting
imise problems associated with the high temperature and high stresses at
the cutting edge of the tool during machining. The main functions of cutting fluids are
lubrication, cooling and less important, to clear the swarf away from the cutting area. It is
fluids are used to min
87
important to understand how cutting fluids work as lubricants as well as coolants. As a
lubricant the cutting fluid works to reduce friction between the tool and the workpiece
material, consequently reducing the seizure region. Therefore, its effectiveness depends on its
ability of penetrating the chip-tool interface and to form a thin layer in the shortest available
time, either by chemical attack or physical adsorption, with lower shear strength than the
strength of the material in the interface. In fact, lubricant can only be effective in the sliding
zone because the cutting fluid whether in liquid or gaseous form, is unable to gain access to
the seizure zone (MACHADO; WALLBANK, 1994). As a coolant the cutting fluids reduces
the temperature generated at the tool-workpiece and tool-chip interfaces both by its cooling
action and by reducing the heat generated during machining (MACHADO (1990), SALES;
DINIZ; MACHADO (2001), BONNEY (2004)). Despite attempts to eliminate the use of
cutting fluids in machining operations, when properly applied, they can either increase the
rate of production or reduce the total cost per part by ensuring possible the use of higher
cutting speeds, higher feed rates and bigger depths of cut (KATO; YAMAGUSHI;
YAMADA, 1972). Effective application of cutting fluids can also prolong tool life, thereby
reducing the number of tool changes, increase dimensional accuracy as well as improve
surface of the machined components (DA SILVA et al., 2001) and decrease the amount of
power consumed. The coolant is drawn into the tool-chip and tool-workpiece interfaces by
capillary action of the interlocking network of surface asperities. The size of these asperities
requires that cutting fluids should have smaller molecular size and good surface tension
characteristics. The boundary lubrication at the tool-workpiece interface reduces the welding
tendency, thus reducing the rate of wear of the cutting tool. Temperature reduction also leads
to a decrease in tool wear rate. This occurs because, first, the tool material is harder and so
more resistant to abrasion wear at lower temperatures, and secondly, the diffusion rate of
constituents in the tool material is less at lower temperatures. Opposing these two effects, a
reduction in the temperature of the workpiece will increase its shear flow stress so that the
cutting force and power consumption may be increased to some extent. Under certain
conditions this can lead to a decrease in tool life. Built-up-edge can be reduced or completely
eliminated by machining with cutting fluids (EL BARADIE, 1996). Since the dual role of
cutting fluids is reduction of the cutting temperature (cooling action) and reduction of the heat
generated by friction, the use of cutting fluids can lead to reduction of cracks in the cutting
tool as well as reduce plucking of tool particles during machining (MACHADO, 1990). The
cutting fluid can intermittently penetrate the flank and rake faces of the tool thereby inducing
88
temperature fluctuation during machining. Due to the elevated cutting temperatures and
periodic penetration of cutting fluid to the cutting interfaces, the cutting tool can be subjected
to continuous expansion and contraction during machining. Therefore, a cutting tool with a
good thermal conductivity and low thermal expansion coefficient can minimise thermal
damages by minimising temperature fluctuation at the cutting edge of the tool (EZUGWU;
BONNEY; YAMANE, 2003).
When choosing a cutting fluid it is important to optimise the beneficial effects and
minimise the harmful effects. In this context, the following properties are desirable in a
cutting fluid (MACHADO (1990), DA SILVA et al. (2001), EL BARADIE (1996),
FRAZIER (1974)):
• Good lubricant;
• High heat absorption capacity;
• It must not emit toxic fumes;
• Harmless to the lubricating system of the machine tool;
• Harmless to the workpiece surface;
e a fire hazard of fume due to increased temperature;
t remain chemically and physically stable for a long
r.
2.14.1 la
one of th in the industries. Most of the cutting fluids fall into one of the three
categ s ZIER (1974)):
. Water based cutting fluids:
ls;
• It should be not b
• It must be economical and mus
period;
• It should maintain a moderately mild or non-existing odo
C ssification of Cutting Fluids
There are several ways of classifying cutting fluids and there is no standard to establish
em with
orie listed below (BONNEY (2004), EL BARADIE (1996), FRA
1
a) Water;
b) Emulsions (soluble oil);
c) Chemical solutions (or synthetic fluids);
2. Neat Oils:
a) Mineral oils;
b) Fatty oi
c) Composed oils;
d) Extreme pressure oils (EP);
89
e) Multiple use oils.
3.
Wat s olant and it is expected to give the best performance due to its high
specific heat, high latent heat of vaporisation and greater turbulence (NAGPAL; SHARMA,
1973). It is, however, a poor lubricant. Because of its high corrosion ability in ferrous
materials, s as a cutting fluid.
u ls as the emulsions are bi-phase composites of mineral oils, additives
(emul
e the formation of small particles of oil,
which
tors, nitrates for nitrite stabilization, phosphates and borates
for w
Gases
er i the best co
it i practically ignored
Sol ble oi
sifiers) and corrosion inhibitors dispersed in water in proportion that varies from 1:10 to
1:100. Additives are usually soaps, acid sludges, saponified phenol and saponified napthenic
acids which act to decrease the surface tension forming a stable monomolecular layer in the
oil-water interface. Therefore these additives provid
can result in transparent emulsions. The stability of the emulsions is related to the
development of an electrical layer in the oil-water interface. Repulsive forces among particles
of the same charge avoid their coalescence. To avoid the bad effects from water of these
emulsions, anticorrosive additives, like sodium nitrite, are used. Since emulsifiers are
susceptible to bacterial attack, biocides are also in the formulation to discourage bacterial
growth. They must not be toxic and harmful to the human skin. Emulsifiable mineral oils have
the desirable characteristics of a good rust inhibition and adequate lubricity for the ordinary
cutting applications. The EP and antiwear additives that increase the lubrication properties are
the same used in neat oils. However, the use of chlorine in cutting fluids has being restricted
due to the harm it causes to the environment and to human health. Sulphur and calcium based
additives are being used instead. Both animal and vegetable grease can also be used to
enhance lubrication properties.
Cutting fluids can also be produced by chemical (synthetic) and semi-chemical (semi-
synthetic) solutions. Synthetic fluids do not have oil in their composition. They are chemical
solutions consisting of inorganic and/or other materials dissolved in water and containing no
mineral oil. They are based on chemical substances that go into these fluids including amines
and nitrites for corrosion inhibi
ater softening, soaps and wetting agents for lubrication and reduction of surface tension,
phosphorus, chlorine and sulphur compounds for chemical lubrication, glycols as blending
agents and humectants, and germicides to control the growth of bacteria. The semi-synthetic
fluids also known as microemulsions are essentially a combination of chemical fluids and
emulsifiable oils in water. These fluids are actually performed chemical emulsions that
90
e
load carrying properties of
lubricant in m
contain only a small amount of emulsified mineral oil, about 5% to 30% of the base fluid,
which has been added o form a translucent, stable emulsion of small droplet size. They are
considered cleaner, with better rust and rancidity control than emusifialbe oils. Extreme
Pressure (EP) and anticorrosion additives can be incorporated like in the soluble oils and thus
making them suitable for moderate and heavy duty machining and grinding operations.
Neat oils are classified into mineral oils, fatty oils and a mixture based on mineral oil
and fatty oil or also with additives, generally of extreme pressure (EP) type. The mixture of
mineral and fatty oils is formulated by blending straight mineral oil with 10% to 40% fatty
oils. The amount of fatty oils used is dependent on the machinability of the material
concerned. Additives are added to neat cutting oils because in some machining operations th
straight oil are inadequate for the severe conditions experienced in
the cutting zone. Small additions of the fatty oils have the effect of markedly improving anti-
friction characteristics under conditions of boundary lubrication when the rubbing faces are so
heavily loaded that a straight oil would be unable to keep the faces apart the fatty additives
forms a thin and highly tenacious layer of metallic soap, created by chemical activity between
the fatty acid molecules and the metal of the tool, the chip and the workpiece. This layer has
very low shear strength and continues to lubricate even after the normal oil film has broken
down. The use of these oils as cutting fluids has decreased due to the high cost, fire risks,
inefficiency in high cutting speed, low cooling ability, smoke formation and high risk to the
human health when compared to water based cutting fluids. Extreme pressure (EP) additives
are added to cutting fluids used for machining operations where cutting forces are particularly
high, such as tapping and broaching, or for operations performed with heavy feeds. EP
additives provide a tougher, more stable form of lubrication at the chip-tool interface. These
additives include sulphur, chlorine or phosphorus compounds that react at high temperatures
in the cutting zones to form metallic sulphates, chlorides and phosphides.
Gaseous lubricants appear very attractive when the cutting fluid penetration problem is
considered. Operations which are performed dry are actually carried out using air as cutting
fluid because of its very effective boundary lubrication property. However, as a coolant air is
not recommended for machining applications because all gases have relatively poor cooling
ability compared to liquids (EL BARADIE, 1996). Detailed information about gaseous
achining will be given in Section 2.15.6.
91
Table 2.6 - Tool materials properties and cost (ABRÃO, 1995).
al
High speed steel
(M2)
Cemented
carbide
(M20)
White
alumina
Mixed
alumina
Whisker
reinforced
alumina
Silicon
nitride-based
CBN/PCBN Natural
diamond
PCD
Tool materi
Property
Typical
composition*
0.85wt%C
4wt%Cr
5wt%Mo
6.5wt%W
2wt%V
89.5wt%WC
10wt%Co
0.5wt%other
90-95%
Al
2
O
3
5-10%ZrO
2
Al
2
O
3
30% TiC
5-10%ZrO
2
75% Al
2
O
3
25% SiC
77% Si
3
N
4
13% Al
2
O
3
10% Y
2
O
3
98% CBN
2%AlB
2
/AlN
-
PCD
2-8% Co
Density (g cm
-3
)
7.85 14.5 3.8 - 4.0 4.3 3.7 3.2 3.1 3.5 3.4
Hardness at RT
(HV)
850 1600 1700 1900 2000 1600 4000 10000
8000-
10000
Hardness at 1000ºC
(HV)
n.a.
≈400
650 800 900 900
≈1800
n.a. n.a.
Fracture toughness
(Mpa m
0.5
)
17 13 1.9 2 8 6 10 3.4 7.9
Thermal
conductivity (Wm
-1
ºC
-1
)
37 85 8 - 10 12 - 18 32 23 100 900 560
Young´s modulus
(kN mm
-2
)
250 580 380 420 390 300 680 964 841
Coefficient of
thermal expansion
(x10
-6
K
-1
)
12 5.5 8.5 8 6.4 3.2 4.9 1.5 - 4.8 3.8
Approximate cost
per edge
♠
(^)
40.3
(bar 25x25
x200 mm)
0.34 0.46 0.6 2.5 1.25 40-60 125-140 30-50
* by volume unless otherwise stated
♠
cost refers to ISO SNGN 120416 inserts
92
2.14.2 Directions of Application of Cutting Fluids
Cutting fluids are generally directed to the cutting regions where heat is generated
though nozzle (external) com d ere
and the chip, between the tool and the chip, between the workpiece and tool, under flooding
or “overhead flood cooling” and through the tool (internal) gur 2. Which way that is
best will depend on the cutting process and workpiece material. Problems with nozzle holders
and positioning o et have to be considered if an external jet is used. Factors such as tool
layout and tool iting ac us f an external nozzle
(DAHLMAN, 2000). The internal application of the jet is suitable for some cutting processes
e.g. drilling and milling. The application of coolant can cause a change in the distribution and
locatio e p eratur gio n to h han is d tly linked to the effect
of coolant on chip curl and tool wear (DAHLMAN (2000), SEAH; LI; LEE (1995)). Shaw
(1984) found that when the flow is directed from A it can affect the chip curl and chip tool
contac s th cation of the point of ma um mp re on the rake face of
the tool. If the point of maximum temperature is brought close to the cutting edge the effect of
applying the flow direction A will be detrimental. Pigott and Colwell (1952) found that tool
wear was reduced when very high velocity jet of id ap ed t gh direction C. Smart
and Tr 974) also carried out a study on the luence of cutting fluids and the directions
luid application on emperature distribution in the cutting tool when machining iron and
el epo ly c n om re n as most effective of all
ctions
bined with iff nt directions such as between the workpiece
, Fi e 2.4
f th
pa
e j
th have severe lim
imp t on the e o
n of th eak temp e re n i the ol. T is c ge irec
t length and thu e lo xim te eratu
flu
inf
was pli hrou
ent (1
of f
nick
dire
t
that
and r
.
rted app ing oola t fr di ctio C w the
93
technique in milling stainless steel. They reported that longer tool life, reduction in
cutting forces and better chip form were obtained when the jet was applied through an
extern zzle toward the area between the rake face and the chip (B direction in Figure
2.42) compared to the jet directed through a hole in the tool rake face (position E) and
overhead (position A). Outstanding performance in drilling of Ti-6Al-4V alloy with high-
pressure internal cooling was obtained compared with external cooling (LÓPEZ DE
LACALLE et al., 2000). Dahlman (2000) found that external application of the jet is suitable
for turning operations whereas the jet applied through internal channels in the tool can be
more favourable for milling operations. A recent work employing different directions (A, B
and C) of application of cutting fluid and different tool geometry in grooving of titanium alloy
was carried out by Norihiko and Akio (1998). They found that by applying the cutting fluid
from B and C directions simultaneously the grooving performance was improved with the
prevention of the adhesion between chip and rake face of the tool as well as the great
Figure 2.42 - Schematic illustration of the possible directions of application of cutting fluids.
Kovacevic; Cherukuthota and Mazurkiewicz (1995) carried out a study to evaluate the
performance of three different methods of application of waterjet to the tool-chip interface
(through a hole in the tool rake face) through an external nozzle and conventional cooling
(overhead)
A – Workpiece-chip
B – Tool-chip
C – Workpiece-tool
D – Overhead
E – Internal channel in the tool
Workpiece
Tool
Cutting direction
Chip
A
D
B
E
C
al no
94
reduction of the adhesion on the tool side flank. Conversely, Seah; Li and Lee (1995) found
that tool life drastically decreased by application of water-soluble cutting fluid (flowrate of
2.5-3 l min
-1
)
from A direction ) relative t ry condition when turning AISI 1045
and AISI 4340 with uncoated cemented carbides at cutting speeds up to 190 m min
-1
. The
authors also reported that one effect of the coolant was to slightly increase the crater wear and
to shift the position of the crater wear nearer to the tool tip. This causes the tool cutting point
to become much weaker.
2.15 Cutting Environments and Techniques Employed when Machining Titanium Alloys
Increased productivity in the manufacturing industry can be achieved by reduction in
both non-machining and/or machining time. It is estimated that the cost of an inactive
mach e cost as that of machine used in production. Prevention of this idle
time (LI; LOW (1994), DAHLMAN (2000)). One common
reason for unwanted idle time is the machine operator stopping the machine to clear the chips
gener
(Figure 2.42 o d
ine is almost the sam
must therefore be given priority
away as well as to index the cutting tool. The in-put variables of the machining system which
ally have influence on machinability of titanium alloys are cutting speed, feed rate,
depth of cut, cutting tool and tool geometry, machining environment and the machine tool
itself. Tool life and chip shape depends on the in-put variables of the machining system.
Reduction of the idle time will depend on a combination of the best machining parameters.
However, this combination is not easy to achieve, especially when machining titanium-alloys
with characteristics that impair machinability such as low thermal conductivity that
consequently increases temperature generated at the tool-workpiece interface, adversely
affecting tool life. Since the main in-put variables (cutting speed, feed rate, depth of cut and
cutting tool) for a particular machining operation have been chosen previously, the correct
choice of cutting environment or atmosphere as well as the suitable technique for machining
any material is a decisive factor in order to achieve higher production rates and/or lower
manufacturing cost (s).
Attempts have been made to control the high temperature at the cutting zone when
machining titanium alloys with cutting fluids using conventional coolant flow. High pressure
coolant delivered to the tool cutting edge can provide adequate cooling at the tool-workpiece
interface and ensure effective chip segmentation and removal chip from the cutting area
95
during machining. Other cooling technique, like the minimum quantity lubrication (MQL) has
shown considerably improvement in the machinability of aerospace alloys compared to
conventional coolant flow (BRINKSMEIER et al., 1999). Although there are still few
information about the use of MQL technique in machining titanium-alloys, some encouraging
results have been reported in comparison to conventional coolant technique. Although
increase in productivity has been obtained using high pressure coolant delivery relative to
conventional methods of co
olant delivery when machining titanium and nickel-base alloys at
latively lower speed conditions, other environments such as atmospheric air (dry
achining), argon enriched environment and liquid nitrogen (cryogenic machining) have been
ue to ecological and economic considerations of the use of cutting fluids. Other
enviro
ls with improved properties, by modification/adaptation of tool
geom
re
m
developed as alternative cooling technology to improve the machinability of titanium-alloys
d
nments such as dry air, oxygen, nitrogen, CO
2
and organic compounds such as
tetrachloromethane (CCl
4
) and ethanol vapour (C
2
H
5
OH) are also expected to improve the
machinability of titanium-alloys by improving the characteristics of the tribological processes
present at the tool-workpiece interface and at the same time eliminate environmental damages
as well as minimizing some serious problems regarding the health and safety of operators
(DA SILVA; BIANCHI (2000), SOKOVIC; MIJANOVIC (2001)).
Despite the poor machinability of superalloys, especially titanium and nickel-base
alloys, some special machining techniques including specially designed ledge tools, self-
propelled rotary tool (SPRT), ramping technique (taper turning) and hot machining have
shown remarkable success in their machining. All of these machining techniques have been
used to retard tool wear and consequently prolonging tool life when machining superalloys.
As temperature plays a major part in tool failure during machining, it would be reasonable to
minimise or even eliminating the temperature generated at the tool-workpiece and tool-chip
interfaces. However, for some reasons, the special techniques seem to be not popularised.
2.15.1 Dry Machining
Research activities have mainly concentrated on optimising manufacturing technologies
by development of too
etries as well as optimisation of cutting conditions in order to enhance machining
productivity (TONSHOFF et al., 1998). In addition to that, ecological regulations and
economical considerations have emphasised the need for more dry machining or
environmentally clean metal cutting processes. Despite the use of cutting fluid in nearly all
96
machining, the technology of dry machining is still being applied for special machining
operations. The handling, maintenance, recycling and disposal of coolant usually involve
significant costs.
In a machining process, new surfaces are cleaved from the workpiece through the
removal of material in the form of chips which demands a large consumption of energy. The
mechanical energy necessary for the machining operation is transformed into heat. As a result
high temperatures, pressures and severe thermal/frictional actions occur at the tool edge in the
cutting zone. The greater the energy consumption, the more severe the thermal/frictional
actions, consequently accelerating tool wear and making the metal cutting process more
inefficient in terms of tool life, dimensional accuracy and material removal rate
(KOVACEVIC; CHERUKUTHOTA; MAZURKIEWICZ, 1995). Therefore, the efficiency of
the metal cutting process depends to a large extent on the effectiveness of the tribological
condition provided for specific material-cutting tool interaction. Unfortunately, conventional
cutting fluids cause environmental and health problems. Process-generated pollution in
machining has been mainly coming from waste cutting fluids. The cutting fluids disposal
became important environmental and economical issues to be considered by machining
industry because of the high cost of disposal. The current attenti
on to the environmental
impac
loy
comp ed to minimum quantity of lubrication (MQL) technique and conventional flooding at
min
-1
, feed per tooth f
z
= 0.08 mm, width of cut a
p
= 5 mm and depth
of cut
ts of machining processes by government regulations has been forcing manufacturers to
reduce or eliminate the amount of wastes (ÇAKIR; KIYAK; ALTAN, 2004). Therefore, dry
machining has become a reliable choice in machining of some materials. Decision for dry
machining must therefore be made on a case-by-case basis. It is known that these factors are
more pronounced especially in dry machining of titanium alloys, thus more wear-resistant
cutting tool materials that possess higher hot hardness as well as powerful and rigid machine
tools are required (GRAHAM, 2000). The lowest tool lives were obtained with high-speed
steel (HSS) and uncoated carbide tools in dry milling of titanium-base, Ti-6Al-4V, al
ar
a cutting speed of 210 m
a
e
= 2 mm (BRINKSMEIER et al., 1999).
Some researches have shown that dry machining can be applied in some materials with
coated tools (TONSHOFF et al. (1998), NABHANI (2001a)). Coatings, generally with
thickness between 2 and 18 µm, can control temperature fluctuation by inhibiting heat transfer
from the cutting zone to the insert or tool. Multilayer PVD coatings with few nanometers
thick also have been used (GRAHAM, 2000). The coating acts as a heat barrier, because it has
97
a much lower thermal conductivity than the tool substrate and the workpiece material. Coated
cutting tools, therefore, absorb less heat and can tolerate higher cutting temperatures,
permitting the use
of more aggressive cutting parameters in both turning and milling without
sacrif
re also recommendable to this application
becau
icing tool life. PVD coating of titanium aluminium nitride (TiAlN) are harder, thermally
stable and chemically more wear-resistant than TiN coating and have performed better in dry
machining at higher speed condition when machining difficult-to-machine materials
(GRAHAM, 2000). The better performance of TiAlN coating it attributed to the amorphous
aluminium-oxide film that forms at the chip-tool interface when some of the aluminium at the
coating surface oxidise at high temperatures. A recent study on the performance of oxygen-
rich TiAlON coating of WC-carbides in dry drilling of tempered steel showed that the best
wear resistance was found for seven-layered films with alternating interlayers of TiAlN and
TiAlON (TONSHOFF et al., 1998). There was no beneficial effect of TiC/TiC-N/TiN coating
deposited in tungsten carbide substrate inserts in dry machining of titanium-base, Ti-4Al-8V,
alloy at a surface speed of 75 m min
-1
, a feed rate of 0.25 mm rev
-1
and a depth of cut of
1.0 mm (ZOYA; KRISHNAMURTHY, 2000). The coating layers were rapidly eroded by a
process of adhesive wear, leaving the tungsten carbide substrate vulnerable to cratering.
Deformed carbide particles were also observed to be carried off by the underside of the chip,
demonstrating the intimacy and strength of the bond between the workpiece and the tool.
CVD coating process can be used for depositing a layer of aluminium oxide, the most heat
and oxidation resistant coating known. Aluminium oxide (Al
2
O
3
) is a poor heat conductor and
insulates the tool from the heat generated during chip formation, forcing it to flow into the
chip. Al
2
O
3
is therefore an excellent CVD coating for most dry turning operations performed
with carbide substrate. It also protects the substrate at high cutting speeds and is regarded as
the best coating material in terms of abrasion and crater wear resistance.
Dry machining of hardened steels has been performed very well using polycrystalline
cubic boron nitride (PCBN) and ceramic cutting tools. It is believed that concentrating heat in
the cutting zone will reduce the workpiece shear strength, therefore, lowering cutting forces.
Pure alumina oxide ceramics (Al
2
O
3
+ ZrO
2
) a
se of their lower thermal conductivity (ÁVILA; ABRÃO, 2001).
There have been many attempts on dry machining of titanium alloys with conventional
and advanced cutting tool materials. These studies have so far provided little improvement in
terms of tool life and surface integrity of machined surfaces. However, recent publications
have reported the possibility of dry machining of hard materials with ultra-hard cutting tools
98
such as PCD and CBN. The use of ultra-hard cutting tools for prolonging tool life when
machining titanium alloys at relatively high cutting speeds has not been required because
titanium alloys are believed to readily ignite at these conditions. PCBN and PCD cutting tools
have, however, been used to machine titanium alloys at lower speed conditions (up to 70
m/min). Owing of their high hardness and high melting point, CBN tools can withstand the
heat and pressure developed during cutting. According to Seco Tools (2002a), surface
integrity of components machined can generally be maintained due to the ability of CBN tool
to maintain a sharp edge even when it has witnessed significant wear. Nabhani (2001a)
reported satisfactory tool life and surface finish when dry machining titanium-base, Ti-4Al-
8V, alloy with ultra-hard materials (PCD, CBN) and coated carbide tools at a surface speed of
75 m min
-1
, a feed rate of 0.25 mm rev
-1
and a depth of cut of 1.0 mm. Average flank wear
was the predominant failure mode. Ultra-hard materials are suitable for finish machining of
titanium alloys at a cutting speed of about 150 m min
-1
and a depth of cut over 0.5 mm (SECO
TOOLS, 2002a). According to this study, PCD tools performed better than CBN and
cemented carbide tools, in terms of wear rate and with over 250% improvement in tool life,
when dry turning Ti-base, Ti-6Al-4V, alloy at a speed of 75 m min
-1
, a feed rate of 0.25 mm
rev
-1
and a depth of cut of 1.0 mm. These results show that ultra-hard tool materials can be
used in dry machining of titanium alloys at lower cutting speeds and at relatively lower feed
rates. Despite the encouraging performance of ultra-hard tool materials, their application in
dry machining of aerospace alloys is still questionable due to their higher cost compared to
carbide tools. Improvement in recorded tool life is not sufficient to compensate for their high
cost, consequently their use in dry machining is not economical. For instance, the price per
cutting edge of a PCD insert is about 22.5 times higher than a carbide insert (SECO TOOLS,
2002b).
Besides titanium alloys, materials like austenitic stainless steel, high-temperature alloys
like nickel alloys and hardened steel demand alternative machining environments
(DALHMAN; KAMINSKI, 1999). The most commonly used environments for machining
aerospace alloys are with conventional coolant supply, high and ultra high pressure coolant
supplies, which generally delivers coolants in variable quantities. Others
include atmospheres
conta
ining nitrogen, CO
2
, argon, ethanol vapour and pure oxygen can also be used singly or in
combination with high pressure coolant delivery systems in machining applications.
99
2.15.2 Conventional Coolant Supply
Conventional coolant supply technology is the most employed technique to deliver
cutting fluid to the cutting zone during machining, especially when machining titanium alloys,
despite the trend towards dry cutting and “ecological” (minimum quantity of lubrication)
machining. Other terms such as flooding, conventional cooling and flood application can also
be encountered in the literature to refer to the conventional coolant supply technique. The
cutting fluid is pumped through a flexible pipe to a nozzle, which can be adjusted to direct the
stream of the fluid to the desired point. The flow is supplied with little (< 0.3 MPa) or no
static pressure but at abundant flow rate (generally lower than 10 l min
-1
)
to ensure complete
coverage of the cutting area (SECO TOOLS, 2002b). This can be achieved by using large
piping with a relief valve in the system and by removing restrictions to get maximum nozzle
flow at low pressure (MACHADO, 1990). Typical directions of application of cutting fluids
when using conventional method are through directions A (workpiece-chip interface), B (tool-
chip interface), D (flooding or “overhead”) as illustrated in Figure 2.42. Conventional coolant
flow technique is effective when machining at lower speed conditions when temperatures at
the cutting zone are relatively low. At higher speed conditions, the cutting fluid has negligible
access to the tool-workpiece or the tool-chip interfaces which are under seizure condition. The
high temperature generated close to the tool edge during machining generally causes
vaporisation of the cutting fluid (BONNEY, 2004).
When using cutting fluids in machining, the coolant flow rate should also be taken into
consideration. It is believed that by increasing the flow rate and the heat transfer rate heat
dissipation at the chip-tool-workpiece interfaces also improves as more quantity of fluid is
delivered to the cutting area. Tool life increased by up to 2 folds when the flow rate of
emulsion
oil was increased from 1.7 to 6.8 l min
-1
during turning of titanium-base, Ti-6Al-4V,
alloy
(MANTLE; ASPINWALL, 1998). However, in some cases, this cooling effect does not
represent advantage in machining. As the heat generated during the machining is transferred
to the workpiece, tool, chip and cutting fluid, more attention should be focused on the cutting
tool. Subjecting the workpiece to continuous cooling will mean that more energy is required
for material removal. High cutting forces therefore becomes inevitable thus a more heat
resistant cutting tool is required.
100
2.15.3 High Pressure and Ultra High Pressure Coolant Supplies
An interesting alternative to conventional coolant supply is the high pressure supply
technique. High pressure coolant supply which delivers coolant to the cutting environment has
been used for some time to improve the machinability of materials, particularly aerospace
alloys. Its credibility had been investigated over the last 50 years and, although practically all
high pressure systems tested have produced benefits to the machining processes, they were
not adopted as a commercial process in the past because of equipment cost and also the fact
that low speed machining was the preferred mode of production as machine tools were not
capable of high speed machining applications (BONNEY, 2004). Today higher production
rate is the basic principle in industry and the most economic means of achieving this is
through the use of high pressure coolant technique. The term “ultra high pressure” is currently
used for coolant pressures in excess of 30 MPa. Cutting fluids can be delivered under high
pressure supply to very close to the critical point on the secondary shear zone (B and E
directions – Figure 2.42 as well as on the workpiece-tool interface (C direction). Figure 2.43
shows a schematic illustration of a toolholder used to deliver coolant at high pressure through
an internal channel that reaches the cutting tool in the B direction (SECO TOOLS, 2002b).
Figure 2.43 - Schematic illustration of a tool holder used for machining under high pressure
coolant supply (SECO TOOLS, 2002b).
101
In addition to ensuring cooling at high speed conditions, the coolant under high pressure
can a
e
middle point of the cutting width and it is assumed that maximum pressure peak occurs just at
the jet hit position. At longitudinal turning when the nose radius is engaged, the pressure build
up will differ from orthogonal turning. The pressure distribution form cannot be Gaussian
since the workpiece helps sustaining a pressure level at the side (DAHLMAN, 2000), see
Figure 2.44 (b). The shape and size of the pressure distribution enable control of the chip form
and flow direction (DAHLMAN (2000), DALHMAN; KAMINSKI (1999), KAMINSKI;
DAHLMAN (1999)). Additionally, the temperature gradient is reduced by penetration of the
high-energy jet into the tool-chip interface and consequently eliminating the seizure effect
(MAZURKIEWICZ; KUBALA; CHOW, 1989), thereby providing adequate lubrication at the
tool-chip interface with a significant reduction in friction (EZUGWU; BONNEY; YAMANE,
2003). These combined with high velocity coolant flow causes the breakage of the chips into
very small segments. Because the tool-chip contact time is shorter, the tool is less susceptible
to dissolution wear caused by chemical reaction with newly generated chips, especially
titanium-alloy chips (LINDEKE; SCHOENIG; KHAN, 1991).
lso act as a chip-breaker. Producing snarled chips during machining operations may
cause problems such as exposing the operator to imminent danger and damage to the
workpiece. This can also hinder the use of advanced manufacturing techniques (unattended
machining, for instance) as well as hindering access of the cutting fluid to the cutting zone
and associated disposal problems that alternately increase machine tool downtime during
production. The high/ultra high pressure coolant jet, generally within the range
0.5 – 360 MPa, is directed via a specially designed orifice in the tool-holder to the region
where the chip breaks contact with the tool. A hydraulic wedge is created at the tool-
workpiece interface which allows the coolant jet to penetrate the interface deeply with a speed
beyond that necessary even for very high speed machining. This action reduces the tool-chip
contact length/area and also changes the chip flow direction. A resulting force acting upon the
chip is created from the pressure in the fluid wedge. The cantilever effect on the chip is
dependant on the pressure distribution. The position of the resulting force can be changed by
altering the jet hit position along the cutting width, Figure 2.44 (a) (DAHLMAN, 2000). From
this approximation of the pressure distribution in orthogonal cutting, the jet is directed at th
102
(a)
(b)
Figure 2.44 - Pressure distribution from the jet momentum action on the chip. (a) Cutting in
tube with single straight edge; (b) pressure distribution (2D) at longitudinal turning
(DAHLMAN, 2000).
The efficiency of high pressure coolant technique is strongly dependent upon the jet
pressure levels (DAHLMAN, 2000). Pigott and Colwell (1952) carried out a series of tests
using various jet pressures from 0.17 to 4.14 MPa in machining of SAE 1020, SAE 3150,
AISI 3140 steels. They surprisingly found that when the jet pressure exceeded 2.76 MPa the
tool life started decreasing. This phenomenon was attributed to an optimum jet-pressure or
critical limit as a result of the critical boiling action of the coolant at the tool edge since it was
possible to sweep the tool surface faster by the higher jet speed, thus lowering the rate of
boiling and cutting down heat transfer which presumably may vary with tool/work material
combination as well as cutting conditions. Nagpal and Sharma (1973) carried out various tests
similar to those of Pigott and Colwell in order to investigate the optimum jet pressure varying
from 0.34 to 3.4 MPa when machining a medium carbon steel with four different cutting
fluids. The pressure of 1.34 MPa was the optimum jet pressure achieved by these researchers
when using three different oils and it appeared to have a relationship with the total heat
generated during machining. However, when using soluble oil the optimum jet pressure was
found to be strongly dependent on the feed rate. Kishi et al. (1975) also investigated the
influence of high pressure coolant technique on machinability parameters such as cutting
forces, cutting temperature and surface roughness when machining pure aluminium, nickel-
base, Nimonic 80A, alloy and various grades of steels using four different soluble oils as
utting fluids. The jet was directed from B position (Figure 2.42) and jet pressure varied from
.3 to 3.9 MPa. They reported that cutting forces, cutting temperature and surface finish all
decrease from a zero value of pressure up to a certain value and were either saturated or
c
0
103
reached a critical point and that increased slightly thereafter. Additionally, it was found that
cutting fluid composition as well as their effects on lubrication and cooling were mutually
related and influenced the machinability of all workpiece materials tested with an increase in
jet pressure. The authors explained that the reason why these parameters reach a constant
value and increase again was due to the permeation depth of cutting fluid into the cutting
zone. As the jet pressure increases cutting fluids can be introduced deeply into the tool-chip
interface and thus the effec lubrication and cooling are sufficientl tained. However,
the permeation depth for cutting fluid to be introduced reaches a constant value and it can not
p
he cutting
ably under constant jet permeation depth conditions. Consequently,
cutting force and other parameters are minimised in such jet pressure. When the jet pressure is
increa
ts of y ob
be introduced more deeply than this depth because of high hydrostatic stresses at the tool-chi
interface, even if the jet pressure is increased. The formation of vaporised films in t
zone decreases remark
sed, the amount of the thick vaporised films to give the lubricant effect will reduce
gradually and instead thin vaporised films will exist in the cutting zone between the tool and
the chip. Consequently, as frictional force increases, cutting temperature and cutting force
increases gradually with increase of jet pressure. Recently Kovacevic; Cherukuthota and
Mazurkiewicz (1995) also investigated the performance of two different methods of
application of high pressure water jet cooling/lubrication with jet pressure varying up to
200 MPa in milling stainless steel. In all of the cases the authors observed that when water
pressure reached a certain optimum value, a further increase in water pressure was not found
to be very beneficial in further improving machining performance. The authors suggested that
this could be due to the fact that a high pressure waterjet after penetrating to a certain depth
into the tool-chip interface is not capable of penetrating any deeper, hence overcoming the
high contact pressures at the tool-chip interface. Bonney (2004) carried out a recent study of
high speed machining of Inconel 718 with ceramic and coated carbide tools at cutting speeds
up to 60 m min
-1
and feed rates varying from 0.1 to 0.3 mm rev
-1
under various high pressure
coolant supplies with jet pressures of 11 MPa, 15 MPa and 20.3 MPa. He reported that
15 MPa was the optimum jet pressure for machining with coated carbide tools at speeds lower
than 30 m min
-1
and a feed rate of 0.25 mm rev
-1
, whereas the jet pressure of 11 MPa was the
optimum pressure for machining with the same tool materials in all cutting speed tested and a
feed rate lower than 0.25 mm rev
-1
. From this study it is clear that optimum jet pressure is
interdependent of the cutting parameters employed, especially the feed rate.
104
Another relevant factor that must be considered when using high or ultra high pressure
coolant delivery is the cooling jet power. Tool wear increased with an increase in the
maximum jet power up to 12kW when machining Ti-base, Ti-6Al-4V, alloy with carbide
tools at a speed of 150 m min
-1
, feed rate of 0.18 mm rev
-1
and depth of cut of 1.5 mm. Crater
wear occurred with further increase in the jet power. Chips produced when machining at a jet
power lower than 3kW were not completely serrated (VIGNEAU, 1997). Besides the correct
choice of coolant pressure, other important factors that can improve the efficiency of coolant
delivery at high pressure must be considered such as jet properties and shape of the nozzle.
As the coherence and stability of the jet are associated with chip-control, a reduction in the
jet velocity leads to a decrease in momentum power at the impact position. Therefore, a stable
pressure build-up at the chip-tool interface is necessary for utilising the effect of the jet in the
best possible way. A free jet beam in the air is exposed to drag reduction. The longer the
travelling distance in air the more the jet beam deteriorates. In some applications long
travelling distances is, however, necessary due to design of tools and other external
conditions. It is recommended to keep the distance from the nozzle to the impact position at a
reasonable level, preferably no longer than 50 mm (NAGPAL; SHARMA, 1973).
The shape of the nozzle forming the jet is important for several reasons:
• Microscopic imperfections present in worn nozzles cause turbulence and cavitation
(NAGPAL; SHARMA, 1973), thereby influencing the boundary layer of the jet.
This accelerates deterioration of the jet produced from a worn nozzle than from a
new one. This action tends to increase turbulence, leading to loss of jet coherence;
• The nozzle geometry should be adapted to each cutting application and can depend
on some factors such as pressure, available flow and tooling;
• Wide rectangular nozzles can cover the cutting edge with one jet. The same
coverage can be achieved with more than one round nozzle.
The use of high pressure coolant supply in machining applications was earlier reported
in the 1950s by Pigott and Colwell (1952) and later on by Field (1968). From that time until
the end of 1980´s few studies (NAGPAL; SHARMA (1973), KISHI et al. (1975), SHARMA;
RICE; SALMON (1971), MAZURKIEWICZ; KUBALA; CHOW (1989)) were encountered
in the literature. However, it was through the 90´s decade that number of studies employing
high pressure coolant technique in machining applications was ra
pidly increased
(MACHADO (1990), GETTELMAN (1991), LINDEKE et al. (1991), WERTHEIM et al.
105
(1992
th
use
f 4kW and
12 kW y, compared to
conve o
pressure o
of 60 m
conventio
pressu c
due to the d when machining the
Incon 9
MPa in fa ide tools reported a 2.5
times increase in tool life than when machining with conventional coolant flow (KLOCKE;
FRIT
), KOVACEVIC; CHERUKUTHOTA; MAZURKIEWICZ (1995), NORIHIKO;
VIGNEAU (1997), AKIO (1998), BRINKSMEIER et al. (1999), CRAFOORD et al. (1999),
DALHMAN; KAMINSKI (1999), EZUGWU; WANG; MACHADO (1999), KAMINSKI;
DAHLMAN (1999), SOKOVIC; MIJANOVIC (2001). More recently other many researches
(DAHLMAN (2000), KAMINSKI; ALVELID (2000), LÓPEZ DE LACALLE et al. (2000),
KLOCKE; RAHMAN; SENTHIL KUMAR; CHOUDHURY (2000), FRITSCH;
GERSCHWILER (2002), BONNEY (2004), DAHLMAN; ESCURSELL (2004),
MACHADO et al. (2004), MAGALHÃES; FERREIRA (2004)) have been carried out wi
of high pressure coolant technique in various machining operations. These studies
generally reported increased productivity when compared to the conventional methods of
coolant delivery. Reduction of temperature in the cutting zone, increase in tool life (up to 5
folds), lower cutting forces, less vibration levels as well as better surface integrity and closer
tolerances of the machined components are major benefits of machining under high pressure
coolant supplies. A 3 fold increase in cutting speed have been reported when machining
titanium-base, Ti-6Al-4V, alloy with carbide tools under high pressure water jet assistance
(HPWJA) relative to conventional flooding (VIGNEAU, 1997). Using HPWJA o
, over 3 and 5 folds increase in tool lives were recorded, respectivel
nti nal coolant flow. Over 300% increase in tool life was achieved using a high jet
f 14.5 MPa in machining Ti-6Al-4V alloy with uncoated carbides at a cutting speed
min
-1
, a feed rate of 0.25 mm rev
–1
and a depth of cut of 2.5 mm compared to
nal coolant flow (MACHADO, 1990). The same author also reported that high
re oolant supply showed much success as a chip-breaker and marginal gain in tool life
high rate of notch wear of the ceramic cutting tool employe
el 01. A recent study of the effect of high pressure coolant flow at a pressure of 14
ce milling of Ti-base, Ti-6Al-4V, alloy with cemented carb
SCH; GERSCHWILER, 2002). Studies carried out at 15 MPa and 30 MPa jet pressures
reported about 50% increase in cutting speed when machining titanium alloys at a higher jet
pressure of 30 MPa relative to conventional coolant flow (SECO TOOLS, 2002a). It has also
been reported that machining under high pressure coolant supply did not lead to a significant
reduction in cutting forces despite the small change in the chip-tool contact length
(EZUGWU; PASHBY (1990), CRAFOORD et al. (1999)). A more recent study (BONNEY,
2004), has shown that high jet pressures of 11MPa and 15 MPa gave about 42% and 71%
106
increase in tool lives respectively when machining Inconel 718 with ceramics compared to
conventional coolant at a speed of 250 m min
-1
and a feed rate of 0.2 mm rev
-1
. With coated
carbide tools it was reported that up to 518%, 647% and 740% improvement in tool lives was
achieved using jet pressures of 11 MPa, 15 MPa and 20.3 MPa, respectively under a cutting
speed of 50 m min
-1
and a feed rate of 0.3 mm rev
-1
. Other recent study (MACHADO et al.,
2004) was carried out to evaluate the influence of the use of high jet pressure technique in tool
life of various cutting tools materials when high speed machining of Ti-6Al-4V alloy. It was
reported that up to 3 folds increase in tool life was achieved when a jet pressure of 20.3 MPa
with uncoated cemented carbide inserts at a cutting speed of 110 m min
-1
and feed rate of
0.15 mm rev
-1
compared to conventional coolant supply. When PCD tools were employed at a
cutting speed of 175 m min
-1
, the gain in tool life was about 21 folds.
Another consideration regarding the use of high pressure system is that machine tools
must be built with very efficient leak-proofing in order to contain the coolant in the machine
and it must also have a mist extractor capable of filtering all the fumes and sprays produced in
the machining environment to avoid harming the operator.
2.15.4 Minimum Quantity Lubrication (MQL)
A machining technique called minimum quantity lubrication (MQL) or minimum
quantity cutting fluid application (MQCFA) is a viable alternative to improving the
characteristics of the tribological processes present at the tool-workpiece interface in order to
improve the machinability of materials and at the same time eliminate environmental damages
as well as minimizing some problems associated with the health and safety of operators (DA
SILVA; BIANCHI (2000), LI et al. (2000), MACHADO; WALLBANK (1997), SOKOVIC;
MIJANOVIC (2001)). Using the same principle of high pressure coolant supply, this
technique consists of applying a small amount of the highly efficient coolant/lubricant which
is pulverised in compressed air stream to the cutting zone at a flow rate often below 100 ml h
-
1
(SECO TOOLS (2002a), LEE (1981)) compared to 120000-720000 ml h
-1
(2-12 l min
-1
)
generally employed in conventional coolant flow. This is why this technique is considered a
“clean cutting process”. MQL also contributes to lowering machining costs as low quantity of
cutting fluid is utilised.
Improvement in machining performance can be achieved when sufficient amount of
cutting fluid gain access to the chip-tool interface (GRAHAM, 2000). Machining with MQL
reduces temperature generated at the cutting zone mainly due to the cooling effect of the
107
compressed air jet and partially due to absorption of heat by the cooling and lubrication
functions of cutting fluids. The fine spray of oil during the process of lubrication of the
cutting zone is also vaporised by the high temperature generated. This small quantity of oil,
most of the time, is sufficient to substantially reduce friction and to avoid the adhesion of the
chip on the tool. Cooling by convection itself is rather low. MQL technique just delivers the
lubricant to the tool-chip and tool-workpiece interfaces while conventional coolant system
floods the entire cutting region. In order to achieve best performance of the MQL technique,
the nozzle geometry must be taken in consideration. In a machining process, good atomisation
is necessary to provide adequate wetting of the work area. Due to exposure to aerosols
(generally consisting of hazardous substances), which offers high risks to human health, care
must be taken when using MQL.
There are limited studies on MQL technique for machining titanium alloys. A recent
study on the use of this technique in milling of titanium alloy with uncoated carbide and high-
speed steel (HSS) tools in three different environments (dry condition, conventional coolant
flooding and MQL technique) at a feed per tooth f
z
= 0.08 mm, width of cut a
p
= 5 mm, depth
of cut a
e
= 2 mm and cutting speeds up to 210 m min
-1
(BRINKSMEIER et al., 1999). The
hen machining with uncoated carbide tools.
When
is that it can easily be mounted on to existing machine tools at
relativ
MQL technique gave moderate performance w
HSS tools were used with MQL technique at a cutting speed of 80 m min
-1
, tool lives
increased by about 32% and 800% compared to conventional coolant flooding and dry
condition, respectively. Despite the high and randomised surface roughness values obtained,
encouraging tool life was recorded with the MQL technique at a flow rate of 10 ml h
-1
compared to dry machining and conventional flooding when turning AISI 52100 hardened
steel with a CBN tool at various cutting speeds (NOVASKI; DÖRR, 1999). The use of MQL
technique can also contribute to reduce machining problems such as pressure welding of chips
to the cutting edge, the main cause of tool failure when milling titanium alloys with high
speed steels (HSS) tools. Improved surface finish of the machined components was obtained
when machining with MQL (EZUGWU; BONNEY; YAMANE, 2003). A two fold increase
in tool life was obtained when milling aerospace alloy steels with coated carbide tools using
MQL technique compared to dry machining (DINIZ; FERREIRA; FILHO, 2003). Another
benefit of this technique
ely low cost. Although MQL technique has been promising, its application in general
machining operations of titanium alloys is unknown and still requires further studies.
108
2.15.5 Cryogenic Machining
The past two decades have witnessed an increase in research investigating the effect of
lower temperatures in the thermal treatment of steels, particularly in HSS cutting tools
(SMOL’NIKOV; KOSSOVICH (1980), POPANDOPULO; ZHUKOVA (1980), TSEITLIN
(1980), ZHMUD’ (1980)). Initial tests consist of subjecting the materials to sub-zero
temperatures varying from –100ºC to –80ºC for a period of 1 hour. Currently, these
temperatures can be as low as – 196ºC (DA SILVA; MACHADO; DE SOUZA, 2001). This is
referred to as sub-zero cooling or cryogenic treatment. The main purpose of this treatment is
to enhance alteration of the microstructure of the material such as transforming the austenite
content present in steels into martensite and precipitation of the fine carbide particles,
consequently altering the tribolog
ical mechanisms to improve its machinability. Up to 25%
impro
vement in the performance of M2 HSS tools cryogenically treated has been reported
when drilling AISI 8640 steel at a cutting speed of 40 m min
-1
(DA SILVA; MACHADO; DE
SOUZA, 2001). However, unsatisfactory results were recorded when milling the same
workpiece material with coated M2 HSS tools cryogenically treated. With technological
advances, it is possible to extend the cryogenic principles to the machining of titanium alloys
as another alternative to improving their machinability, using liquid nitrogen and carbon
dioxide (CO
2
) as the main refrigerants.
Liquid nitrogen (LN2) is often used as coolant because of its higher capacity to absorb
heat than liquefied CO
2
, its relatively low cost and also the fact that LN2 does not harm the
environment. Temperatures in the region of 1100ºC are generated when machining Ti-6Al-4V
alloy (KONIG (1979), MOTONISHI et al. (1987)). This promotes the softening of the cutting
tool material and consequently accelerates tool wear. Application of LN2 to the tool cutting
edge leads to rapid increase in the hardness of the tool, thus retarding tool wear and extending
tool life (MAZURKIEWICZ; KUBALA; CHOW, 1989). Therefore, cryogenic machining is
an efficient way of maintaining the temperature at the cutting interface to well below the
softening temperature of the cutting tool material in addition to being environmentally safe
relative to conventional emulsion cooling. This method consists of delivering LN2 via nozzles
directed to the tool rake face to the flank face or simultaneously to both faces as is illustrated
in Figure 2.45 (HONG; DING; JEONG, 2001).
109
Figure 2.45 - Schematic illustration of nozzle orientation for localized LN2 delivery (HONG;
DING; JEONG, 2001).
Figure 2.46 (MAZURKIEWICZ; KUBALA; CHOW, 1989) clearly illustrates the
cryogenic cooling concept using an insert with an obstruction type chip breaker. In this
schematic diagram liquid nitrogen (LN2) is released trough a nozzle between the chip breaker
and the tool rake face. The chip breaker helps to lift the chip to permit LN2 to reach and cool
the tool-chip interface. Unlike conventional flood cooling, the chip does not block the flow of
LN2. Thus, LN2 absorbs the heat, evaporates quickly and forms a fluid/gas cushion in the
tool-chip interface that acts as a lubricant (HONG; DING; JEONG, 2001). As a result of this,
the coefficient of friction is reduced consequently reducing both crater and flank wear rates.
An auxiliary nozzle can be used to deliver LN2 to the tool flank surface to minimise the
appreciable flank wear on the tool when machining titanium alloys. Figure 2.47 shows a
picture of the tool assembly, nozzles and LN2 flowing out of the nozzle. However, there is an
undesirable phenomena associated with LN2 delivery itself which causes formation of ice on
the channel of the jet flow. This can block the coolant circulatory system, adversely affecting
the cryogenic machining process. A recent comparative study of the wear rates of cemented
carbide tools using LN2 and conventional cooling when machining titanium-base, Ti-6Al-4V,
alloy at a cutting speed of 132 m min
-1
, feed rate of 0.2 mm rev
-1
and a depth of cut of 1.0 mm
showed a five fold increase in flank wear for tools subjected to the conventional cooling
(WANG; RAJURKAR, 2000).
110
Figure 2.46 - A schematic representation of the cryogenic cooling concepts
(MAZURKIEWICZ; KUBALA; CHOW, 1989).
Figure 2.47 -The tool assembly, nozzles and LN2 flowing out of the nozzle
(MAZURKIEWICZ; KUBALA; CHOW, 1989).
111
Yankoff developed an equipment by which carbon dioxide gas (CO
2
) combined with
water-based fluid was directed across the cutting insert at the tool-chip interface
(CRAFOORD et al., 1999). Being a combination of high pressure coolant delivery and
cryogenic principles, expanding and cold CO
2
gas jet at temperature of -73ºC cause breakage
of the stringiest and softest steel chips into very small segments. The gas/fluid stream also
kept the cutting insert cool to the touch, even after machining at high speeds and feed rates.
This system was patented under the name of "Flojet" in 1986. Both CO
2
gas and water-based
fluid emerges from separate lines from a small nozzle under the cutting tool. They are then
mixed in the chamber. Further details on this equipment can be obtained elsewhere
(CRAFOORD et al., 1999). A probable question is why they did not use high-pressure fluid
alone? The answer is based on the heat generation during machining. Both the high pressure
system and the cutting process generate heat. During machining, the water-based fluid soon
warms up with subsequent evaporation that reduces its heat absorption capability. The
excellent cooling capability of the CO
2
gas can keep the fluid, cutting tool and workpiece at
aterial therefore
n be achieved and the need for a fluid
frigeration system is eliminated.
2.15.6 Other Atmospheres
It has been reported from the literature that some gases have replaced cutting fluids in
machining when cutting fluid penetration problem is considered as well as to prevent
oxidation of the workpiece and the chips (EL BARADIE, 1996). Various results were given
such as lower cutting forces and better surface finish. It was noted that the general
information was brief and based on very limited studies. Although these studies have been
carried out on the use of gases such as argon, dried air, oxygen, nitrogen, helium, CO
2
and
ethanol vapour as additional methods to improve machinability of aerospace alloys, many
information on the experimental procedures (specification, methodology and equipment) are
still unclarified. It is known that the principles underlying the chemical processes of these
gases can alter the tribological conditions existing between two surfaces in contact
AWAID (1982),
f gases on the friction and wear of metals
ust be considered and extended to machining operations. These gases to a greater extent are
reactive with tool materials and newly generated machined surfaces.
room temperatures or even lower. The thermal expansion of the work m
remains unaffected, hence longer tool life ca
re
(MISHINA, 1992), such as the cutting zone during machining (J
(KHAMSEHZADEH (1991)). Thus the effects o
m
112
Since these principles can be employed to improve machining productivity, many
studies were prompted to investigate the existing wear pattern between different materials
using the pin-on-disk testing apparatus which has already been used to study coatings for
some time. For instance, the friction coefficient of tungsten lowered by hydrogen, while argon
and nitrogen do not affect wear of nickel (MISHINA, 1992). Titanium and iron have a high
chemisorption activity for oxygen but have no (or very little) chemisorption activity for
nitrogen. “Chemisorption” refers to chemical adsorption where the particles stick to the
surface by forming a chemical bond and tend to find sites that maximize their coordination
number with the substrate (ATKINS, 1990). Results of an investigation between the
relationship of pin-on-disk testing apparatus and machining tests with tungsten-carbide-base
tools in the presence of nitrogen jet flow show that tool wear can be reduced by reducing the
oxygen concentration in the atmosphere at the tool-workpiece interface (TENNENHOUSE;
RUNKLE, 1987).
A major factor hindering the machinability of titanium alloys is their tendency to react
with most cutting tool materials, thereby encouraging solution wear during machining.
Machining in an inert environment, such in argon-enriched environment, is envisaged to
minimize chemical reaction at the t
ool-chip and tool-workpiece interfaces when machining
commercially available titanium alloys at higher cutting conditions. Argon has been mainly
tect welds against oxidation, as well as reducing fume emissions
durin
lubricant characteristics
employed in welding to pro
g welding, and in other operations that demand a non-oxidizing atmosphere without
nitrogen. The use of argon in machining environment has not been fully explored as it is a
novel approach with limited investigation (JAWAID, 1982). It was reported that argon tends
to enhance tool wear rate when machining austempered ductile iron with various ceramic
cutting tools in the presence of argon and in air enriched atmosphere (MASUDA et al., 1994).
A more recent study (EZUGWU et al., 2005) was carried out to evaluate of the effect of
argon-enriched environment compared to conventional cooling environment in high speed
turning of titanium-base, Ti-6Al-4V, alloy with uncoated cemented carbide tools. The argon
gas was delivered to the cutting zone (B direction from Figure 2.42) through a hose at a
constant flow rate of 12 l min
-1
.
Higher nose wear rate and, consequently, lower tool life was
achieved in the presence of argon compared to conventional coolant supply. This was
attributed to the poor thermal conductivity of the argon as well as the poor lubrication
characteristics, which tend to concentrate more heat at the cutting region, thus weakening the
strength of the cutting tools and accelerating the tool wear. Its poor
113
was a
encouraging longer tool life (OKEKE, 1999). It has also been reported that
affini
lso responsible for increasing friction at the cutting interfaces during machining and an
increase in cutting forces required for efficient shearing of the workpiece (EZUGWU et al.,
2005).
During machining operation the cutting edge generates a freshly cut surface. It is known
that the newly generated surface is highly reactive and chemically combines with the
atmosphere. An oxide film is rapidly formed on the new surface when machining in an
oxygen environment (considered as an active atmosphere). This phenomenon is referred to as
“gettering” (KHAMSEHZADEH, 1991). A deoxidised surface layer may enhance the
diffusion-attrition wear process as this can result in a surface which generally is very
susceptible to bonding. When machining in the presence of oxygen enriched atmosphere,
enough oxygen will help to suppress potential bonding. Tool materials with high affinity to
oxygen may become oxidised. These phenomena tend to hinder notch formation by lowering
forces locally, thus
ty of the Ti and Cr to oxygen has enabled them to form oxides which subsequently
prevented the erosion of the beta prime crystals when machining nickel-base, Inconel 901,
alloy with SiC whisker reinforced Al
2
O
3
ceramic tools at high speed (150-215 m min
-1
)
conditions in oxygen, nitrogen and argon enriched environments. However, when machining
in presence of argon and nitrogen enriched atmospheres, severe notching and shorter tool
lives were obtained. These can be attributed to the absence of oxygen from the vicinity of the
cut and consequently higher bonding forces leading to easier “pullout”. Other research states
that cutting force behaviour will depend on particular workpiece-cutting tool combination
used in the presence of an oxygen atmosphere at pressures greater than 100 Pa (DOYLE;
HORNE, 1980).
Çakir; Kiyak and Altan (2004) in a recent study of the effects of applications of gases
such as nitrogen, oxygen and carbon dioxide, as well as cutting fluids and dry condition in
turning of AISI 1040 steel with cemented carbides reported that lower cutting forces and
better surface finish were achieved with application of all the gases compared to dry and wet
conditions. Carbon dioxide gas gave the best cooling effect and provided lower friction
coefficient and the lowest cutting force and thrust force than oxygen and nitrogen.
Applications of gases provided higher shear angle values. The reason for better surface finish
in gas applications, instead of wet machining, was due probably to the low penetration and
worse lubrication effects than cooling (ÇAKIR; KIYAK; ALTAN, 2004).
114
CCl
4
has a good lubricant property at low cutting speeds. Ethanoic acid belongs to the
family of organic fatty acids, which generally form successful boundary lubricant with longer
carbon
chains in conventional sliding tests. Its action involves the formation of a layer of low
shear
lubricant, can, however, produce a marginally lower friction coefficient in
the va
roceeds, they first
achieve maximum flank wear and then the length of the overhang wears back without further
strength metal soap. Ethanethiol can be chemically active because of sulphur in its
compositon. The organic alcohols such as ethanol in liquid phase are generally not attractive
as boundary lubricants and are also considered to be inferior cutting lubricants because they
can normally only adsorb physically on to metal surfaces (WAKABAYASHI; WILLIAMS;
HUTCHINGS, 1992). Based on the above properties and the possibility of success of these
gases in machining applications, a recent study (WAKABAYASHI; WILLIAMS;
HUTCHINGS, 1992) was carried out to evaluate the performance of organic compounds as
gaseous lubricants such as amino ethane (C
2
H
5
NH
2
), ethane (C
2
H
6
), ethanol (C
2
H
5
OH),
ethanethiol (C
2
H
5
SH), ethanoic acid (CH
3
COOH) and tetrachloromethane (CCl
4
) as well as
air and vacuum in orthogonal cutting of an aerospace aluminium alloy (2014A). The
effectiveness of gaseous lubricants on the cutting process was measured by the friction
coefficient. The authors reported that with exception of ethane, the presence of any of the
organic gases provided reduction of friction coefficient. The ranking order for effectiveness of
machining environments was as follows: CCl
4
, C
2
H
5
OH, C
2
H
5
SH, C
2
H
5
NH
2
, CH
3
COOH,
C
2
H
6
, vacuum and air. The best performance of vapour tetrachloromethane (CCl
4
)
environment was attributed to its improved lubrication, which decreased the tool-chip contact
length as well as minimised adhesion wear mechanism between the chip and the tool. Higher
friction coefficient and cutting forces were observed when machining in the presence of air
compared to vacuum condition. Ethanol (C
2
H
5
OH), which a liquid phase is not a very
effective cutting
pour phase.
2.15.7 Ledge Cutting Tools
Ledge cutting tools were developed by the General Electric Company (KOMANDURI;
LEE (1984), DEARNLEY; TRENT (1985)) and characterised by a thin cutting edge (0.38 to
1.27 mm) that overhangs a small distance (0.38 to 1.27 mm) equal to the desired depth of cut.
The optimum thickness will depend on the material to be machined, the cutting conditions and
the flank wear tolerable. The advantage of these tools is that they have a restricted clearance
face which will limit the maximum flank wear of the tool. As cutting p
115
devel
opment in flank wear due to a restricted clearance face. These characteristics allow the
tools to perform for a long period of time as tool life is not limited by amount of flank wear
but by the size of the tool edge (KOMANDURI; LEE, 1984). It was reported that a significant
improve in tool life (30 min) was achieved when a ledge tool was employed in turning
titanium-base, Ti-6Al-4V, alloy with inserts prepared from various grades of uncoated
cemented carbides with 0.8 mm overhang x 1 mm thick x 12.7 wide at cutting speed of
600 m min
-1
(KOMANDURI; LEE, 1984). This cutting speed was up to 5 times higher than
conventional cutting speeds for machining titanium alloys. Figure 2.48 shows a ledge tool
schematically. Because of its restricted geometry these tools are applicable only to straight
cuts in turning, facing, boring, face milling and some peripheral milling operations. Another
limitation of the ledge tool is the depth of cut which must be equal to or less than the width of
the ledge overhang.
(after KOMANDURY; LEE, 1984)) Figure 2.48 - Ledge tool
2.15.8 Rotary Tools
The rotary cutting tool concept has a long history. Rotary cutting tools are in the form
circular discs that rotate about their central axis and the principal difference between rotary
cutting and conventional cutting is the movement of the cutting edge in addition to the main
cutting and feed motion (Figure 2.49). The rotation of cutting edge is the main advantage of
rotary cutting tools when compared with traditional cutting tools as is causes a portion of the
116
tool cutting edge to be heated only for a very short time interval followed by a long rest
period which permits the conduction of thermal energy, associated with the cutting process,
away from the cutting zone (WANG; EZUGWU; GUPTA (1998), EZUGWU; OLAJIRE;
WANG (2002)). Tool wear, in this case, is distributed uniformly across the entire round
shaped cutting edge.
Figure 2.49 - Schematic representation of principle of rotary cutting action (WANG;
EZUGWU; GUPTA, 1998).
The cutting tool rotation is affected either by an external driver, in which case the tool
will be known as a driven rotary tool (DRT), or by the self-propelled action (SPRT) of the
cutting forces exerted on the tool through adjustment of its axis at an inclination with respect
to the cutting velocity (WANG; EZUGWU; GUPTA, 1998) . About 60 folds and 4 folds
improvement in tool life were obtained (EZUGWU; OLAJIRE; WANG (2002), LEI; LIU
(2002)) when machining titanium-base, Ti-6Al-4V, and nickel-base, 718, alloys respectively
at high speed conditions with DRT compared to conventional single point turning with round
unc an be attributed to the very
low flank wear rate as a result of reduced amount of work done in deformation and friction at
ne and on the rake face, respectively, as well as the improved heat
transf
oated carbide inserts The outstanding performance of SPRT c
the primary shear zo
er from the cutting zone due to rotation of the tool edge during machining process.
Additionally, SPRT lower cutting forces, up to 20%, than those obtained under conventional
turning with round inserts. Other advantage of using the SPRT relative to the conventional
turning is that machined surface is less affected by the thermal damage with better surface
finish and compressive residual stresses (WANG, 1997). Reduction in inclination angle
117
improved values of surface finish when machining Ti-6Al-4V alloy with the SPRT compared
to conventional turning operation (WANG; EZUGWU; GUPTA, 1998). However, a SPRT
needs to have a large inclination angle for its insert to be able to rotate and its rotational
speeds depends on machining conditions. This disadvantage limits the surface finish it can
achieve and render it
extremely difficult to optimise the process for longer tool life, good
surface finish and high material removal rate (LEI; LIU, 2002). The limitations of this
technique are that it can only be used at low depth of cuts, generation of more errors than a
stationary cutting edge, developing of chatter due to the large tool radius and poor stiffness of
the rotary system and difficulty in producing components with complex geometry. More
details of the rotary cutting tool technique can be obtained elsewhere (ARMAREGO; KARRI;
SMITH (1994), EZUGWU; WANG (1997), WANG (1997), WANG; EZUGWU; GUPTA
(1998), EZUGWU; OLAJIRE; WANG (2002), LEI; LIU (2002), KISHAWY; WILCOX
(2003)).
2.15.9 Ramping Technique
ckel and
velopment of notch wear (in “V” form) at the tool nose and/or
depth of cut region. Notch wear is formed by mechanical and chemical actions between the
cuttin
A typical problem encountered in the machining of super-alloys such as ni
titanium-base alloys is the de
g tool and the work material during machining. As the notch wear generally grows on a
random and unpredictable basis, high stress are concentrated at the cutting edge making it
prone to catastrophic failure with further machining. Ramping (or taper turning) technique
was developed to minimise or even eliminate notching when machining mainly nickel-base
alloys. This technique consists of varying the depth of cut during the machining in order to
distribute the concentration of notch wear along of cutting edge. The depth of cut line is
shifted gradually along the flank face from a minimum to a maximum value and vice-versa,
thus shifting the notch wear along the entire flank face of tool (EZUGWU; BONNEY;
YAMANE, 2003). This leads to the generation of a uniform flank wear on the cutting edge
thus preventing premature tool fracture, initiated by the otherwise severe notching which may
have been generated, thereby improving tool performance (BONNEY, 2004). High speed
machining trials on nickel-base alloys, such as Inconel 718 and Incoloy 901 alloys, showed
that notching on the tool materials was suppressed by employing ramping technique
(BALAZINSKI; LITWIN; FORTIN, 1993). A recent study (BONNEY, 2004) reported that
ramping technique induced high cyclic loading on ceramic tools when turning Inconel 718
118
alloy, which consequently contributed to development of chipping process and in some cases
flaking on the rake face tools. However, the author observed that notching was completely
eliminated in most of cases when machining with coated cemented carbide tools.
2.15.10 Hot Machining / Hybrid Machining
Hot machining process has attracted considerable attention in the machining of difficult-
to-machine materials such as high manganese steel, nickel and titanium-base alloys. These
materials generally possess high tensile strength and resistance to wear which impair their
machinability. Hot machining principle is based on the heating of workpiece material during
machining. It has been reported that machining by softening of the workpiece through heating
process can be more effective than strengthening the cutting tool in terms of improving
machinability of those materials. For machining it is necessary to choose the best method to
heat the m
aterial ideally. The wrong selection of heating method can induce undesirable
ructural changes in the workpiece and increases the cost. In the researches there are several
used for heating the workpiece. Furnace, gas torch, induction coil,
carbo
st
heating methods, which are
n arc, plasma arc and electrical resistance were the mostly employed methods in
previous studies (SHAW, 1984). Some of these methods have still been used and different
heating methods have been introduced in machining recently (MADHAVULU; AHMED
(1994), BARNES; PASHBY; MOK (1995), RICE; SALMON; ADVANI (1996), PRASAD;
VIGNEAU (1997), SESHACHARYULY (1998), ÖZLER; ÍNAN; ÖZEL (2001), TOSUN;
ÖZLER (2002), AMIN; YANTI; EIDA (2003)). CO
2
atmosphere was used in turning and
milling of nickel-base, Inconel 718, alloy with ceramics and carbide tools at a feed rate of
0.18 mm rev
-1
, depth of cut of 1.5 mm and cutting speeds up to 350 m min
-1
(VIGNEAU,
1997). This process was referred to as the laser-assisted machining (LAM), where a CO
2
gas
jet is directed to the cutting edge at 6kW. Laser-assisted machining uses a laser beam focused
on the work material just in front of the cutting tool. The laser beams extends over the entire
depth of cut and is positioned and adjusted to soften the material on the shear plane without
allowing appreciable laser energy to flow into the tool (SHAW, 1984). This technique
reduced tool wear rate and cutting forces by about 30% and 20-50% respectively when
machining with ceramic tools, while accelerated tool wear rate was achieved with carbides
tools. Similar results limited the use of LAM when machining titanium-base, Ti-6Al-4V, alloy
with carbide tools (VIGNEAU, 1997). Barnes; Pashby and Mok (1995) also employed the
LAM technique and conventional machining during turning of an aluminium/silicon carbide
119
MMC with coated cemented carbides at cutting speeds up to 90 m min
-1
. The material was
preheated at temperatures in the range of 200ºC-400ºC. Tool wear rate increased with
increasing workpiece temperature. The main reason for the detrimental effect of
pre-heating
chnique on the MMC machinability was found to be a shift in the stability range of the built-
rease in flank wear with increasing pre-heat
tempe
te
up-edge to lower cutting speeds. The inc
rature was attributed to a reduction in the protection provided to the flank face by the
built-up-edge. On the other hand, pre-heating the workpiece was beneficial in reducing both
the cutting forces and shear yield stress on the primary shear plane. Drawbacks of laser beam
heating method are regard its low efficiency and high cost as very high power lasers are
required.
The process of plasma-assisted hot machining process for turning operations utilises a
high temperature plasma arc in the range of 16,000ºC to 30,000ºC to provide a controlled
source of localised heat which softens only that small portion of the work material removed
by the cutting tool in the form of chip retaining the metallurgical features of the remaining
portions (MADHAVULU; AHMED, 1994). A plasma arc consists of a high velocity, high
temperature stream of ionised gas capable of supporting a high-current, low-voltage electric
arc. A plasma torch produces this phenomenon by having a tungsten electrode centrally
placed within a water cooled copper nozzle. A gas stream is fed down the annulus between
these, the gas being ionised by a high frequency discharge between the cooper nozzle and the
central electrode. This is followed by a low-current pilot arc and then by a high-current main
arc. The arc characteristics and reliability of arc striking are improved with the balanced
geometry of the nozzle orifice. Madhavulu and Ahmed (1994) carried out a comparative study
of a plasma-assisted hot machining and conventional process in turning various grades of
steels with cemented carbides. They reported that up 2 folds improvement in tool life and 1.8
times gain in metal removal rate were obtained using hot machining process. Additionally
lower power consumption was obtained in the hot machining process relative to conventional
machining process.
Hot machining technique was evaluated by Özler; Ínan and Özel (2001) compared to
conventional technique during machining of austenitic manganese steel. They utilised heating
temperatures ranging from 200 to 600ºC. The flame was generated by a torch containing a
mixture of liquid petroleum gas and oxygen. Figure 2.50 shows the schematic representation
of a hot machining technique design used in their experiments. Longer tool lives were
achieved when the hot machining technique was employed compared to conventional
120
machining. The tool wear rate decreased with increase in heating temperature due to the
reduction in resistance of the workpiece material to shearing. The longest tool lives were
obtained when a heating temperature of 600ºC was used. In other study carried out by Tosun
and Özler (2002) liquid petroleum gas (70% propane and 30% butane) flame was used to heat
the workpiece (heating temperatures ranging from 200 to 600ºC) during a turning operation of
austenitic manganese steel with M20 sintered carbide tools. Artificial neural networks and
regression analysis method were also employed using experimental data to predict tool life.
Tool life increased with increase in heating temperature. 400ºC was the optimum heating
temperature because of the microstructure of the workpiece obtained and the cost.
Figure 2.50 - Schematic representation of a hot machining technique design (ÖZLER; ÍNAN;
ÖZEL, 2001).
Amin; Yanti and Eida (2003) investigated the effect of hot machining process using
oxy-acetylene flame on chatter and tool performance during turning of AISI 1040 steel and
reported that chatter was substantially reduced with lower tool wear rate and better surface
finish using hot machining process. They also reported that diffusion was the predominant
wear mechanism when turning with hot machining process whereas attrition and diffusion
were predominant without preheating.
121
2.16 Surface Integrity
The term surface integrity refers to the control of both surface topography (and
geometry) and metallurgical alteration below the surface obtained after machining with
respect to the base material. In other words surface integrity determines the quality of the
machined surfaces. Temperature generated at the cutting interface during machining generally
induces residual stresses (tensile or compressive), can promote metallurgical alterations,
surface plastic deformation of the work material as well as influence tool wear. Cutting tools
with high wear lands promote severe plastic deformation tearing and cracking of
workpiece
surfaces (BONNEY, 2004). Control of the surface integrity, therefore, is a key factor in
enhancing the function and reliability of a component in order to give the most suitable
condition for long life, fatigue performance, safety of individuals and maximum efficiency at
minimum cost (MACHADO (1990), BONNEY (2004)). The surface integrity is evaluated
either by the surface finish/surface texture and subsurface changes.
2.16.1 Surface Finish and Texture
Surface finish and surface texture are concerned with the geometric irregularities and
the quality of the surface. It consists of roughness, waviness, lays and flaws Figure 2.51
(KALISH (1978), DROZDA; WICK (1983), SCHAFFER (1988), KALPAKJIAN; SCHMID
(2000)).
i) Roughness is the finer random irregularities which usually result from he inherent
t or depth of irregularities is measured in a relative small length called
“roughness sampling length” or “cut-off”. Ideal surface roughness is caused by the
predominant surface patterns and determined by the
production process used. Typical operations that produce pronounced lays are
turning, drilling, milling and grinding;
t
action of the cutting process, as feed marks, instead of from the machine. The mean
heigh
given tool shape and the feed rate;
ii) Waviness is the wider spaced repetitive deviation that may be attributed to the
characteristics of the machine tool, deflection of the workpiece due to cutting loads,
defects in the structure of the work material, vibration, chatter and cutting
temperature, tool wear and to external environmental factors;
iii) Lays denotes the direction of
122
iv) Flaws are characterised as unintentional, unexpected and unwanted interruptions
encountered in the typical topography of a component surface. They are mainly
caused by inherent defects such as inclusions, blowholes, cracks, scratches, nicks
and ridges and also with the presence of the built-up-edge. Flaws can be either
formed during the manufacturing of the material or during the machining of the
material. Unless otherwise specified the effect of flaws is usually not included in the
roughness average measurements.
Figure 2.51 - Standard terminology and symbols of the elements of surface texture (µin)
(KAL K
The
basic cate
i) rmined solely by peak height or valley depth or
ii)
iii) on of amplitude and
The surface roughness is using the arithmetic
average parameter, R
a
(µm), which belongs to the amplitude category. It is obtained by
PA JIAN; SCHMID, 2000).
surface texture is determined by inherent parameters which are classified into three
gories (SCHAFFER, 1988):
Amplitude parameters: they are dete
both, of profile deviation, irrespective of their spacing along the surface;
Spacing parameters: these parameters are determined solely by the spacing of
profile deviation along the surface;
Hybrid parameters: they are determined by the combinati
spacing parameters.
most common method of designating
123
measu
centreline
the centre
by the Eq
ring the mean deviation of the peaks and valleys from the centreline of a trace, the
being established as the line above and below which there is an equal area between
line and the surface trace (SHOUCKRY; 1982), as shown in Figure 2.52 and given
uation (2.13):
n
hhhh
n
++++ ...
321
(2.13)
R
a
=
Figure 2.52 - Schematic illustration of the determination of some amplitude parameters of
surface texture (SHOUCKRY; 1982).
In turning operation, Ra (µm), parameter can be expressed as a function of the feed rate
and nose radius of a tool given by the Equation (2.14) (KALPAKJIAN; SCHMID, 2000):
123
n
h
h
L
r
f
R
a
2
0321.0 ×
=
(2.14)
rev
-1
and r is the tool nose radius in mm. This equation is
only valid when the feed rate is smaller than the nose radius.
observed that increasing the nose radius improves the
surfac i
chatter du increases,
the sh
of the wo . The surface finish can also be improved by
decre
rate and h machining time.
he stylus instrument is the most common instruments employed in measurements the
surfac
Where f is the feed rate in mm
From the equation 2.14 it can be
e f nish. However, too large a nose radius is not recommendable as it tends to promote
ring machining. This may occur because the tool-workpiece contact area
earing stress is decreased, hence higher cutting forces are required for efficient shearing
rkpiece material (BONNEY, 2004)
asing the feed rate. However, reduction in feed rate will decrease the material removal
ence implying longer
T
e finish. The stylus, typically with a diamond tip, traverses across the surface at a
controlled rate and its vertical movement is converted into an electrical signal by a sensitive
124
electronic transducer. Modern instruments digitise the electrical signal and store the results in
a computer, where statistical analysis can be carried out.
2.16.2 Subsurface Chang
es
During the machining process of metal removal is created a surface with a
(extending to a depth of up to 0.25 mm) which can be different from the interior of the
orkpiece. Various subsurface changes can be encountered in machined components. The
bsurface changes consist of several mechanical and metallurgical factors. The main
mechanical factors are (FIELD; KAHLES; CAMMET (1972), DROZDA; WICK (1983)):
• Plastic deformation: it is a common microstructural alteration resulting from
exceeding the yield point of the m nd is normally identified by elongation of
grain structure in the direction of flow and also by the increase in hardness level.
ed,
properties of workpiece materials. Plastic deformation
usually occurs when machining under heavy machining conditions;
ain
attached to the machined surface;
• Microcracking and macrocracking: these are more pronounced when machining
brittle materials (such as cast iron and hardened steel). Microcracks are generally
the region of built-up-edge and
during thermal removal processes;
, shallow or deep. The tensile stress is
layer
w
su
aterial a
The extent of plastic deformation is dependent on the cutting conditions employ
tool geometry as well as the
• Plastically deformed debris: these are generally fragments of a built-up-edge that
has ploughed through the flank surface of the tool and the workpiece which rem
detrimental to fatigue and stress corrosion and thus must be avoided. Cracking can
occur in the vicinity of untempered martensite, in
• Residual stress: it is induced into the surface layer as a consequence of the removal
of machining load that caused plastic deformation on the material as well as intense
localised heating in the surface layer and consequent cooling. The residual stresses
introduced by machining are greatly influenced by the wear land developed on the
tool surface. They also depend on the cutting force, the distribution of temperature
and mechanical stresses during machining operation. Residual stress can be
compressive or tensile, low or high
detrimental to fatigue strength while the compressive stress is beneficial;
• Microhardness alteration: microhardness distribution is dependent on the machining
process and the parameters of that process. The surface hardness can be increased as
125
a result of formation of untempered martensite or plastic deformation by cold work,
i.e., at temperature below the recrystallization t
emperature. On the other hand, the
presence of overtempered martensite will tend to produce surface softening.
ents are a powerful tool used to confirm visual alterations
ed
•
altered.
Microhardness measurem
and help identify changes in machined components, which are not discernible in a
microscopic examination of the surface layer.
In conventional machining the main metallurgical factors are:
• Phase transformation: phase transformation on the surface of the machin
component is caused by the intense heating of the surface layer during machining
process. When machining steel a very brittle and detrimental untempered martensite
can be formed;
Recrystallization: can occur in any metal whose surface suffer excessive
heating and when is plastically deformed during the machining operation. As
a consequence the original properties of the materials can be
CHAPTER III
EXPERIMENTAL PROCEDURE
3.1 Introduction
The objective of this work is to investigate the performance of recently developed
cutting tool materials (coated and uncoated cemented carbides, polycrystalline diamond
(PCD), cubic boron nitride (CBN/PCBN), and ceramics) and various cooling media such as
conventional coolant flow, high pressure coolant supplies (7 MPa, 11 MPa and 20.3 MPa) and
argon enriched environment at cutting speeds up 500 m min
-1
in order to provide a step
increase in the productivity of finish turning of commercially available titanium-base, Ti-6Al-
4V, alloy which is widely used in the aerospace industry. This section describes the
equipment, experimental techniques and methodology employed in this study. All the
experimental tests carried out in this research as well as sample preparation and analysis of
the worn tools and machined surfaces were performed at the laboratories of the Faculty of
Engineering, Science & the Built Environment of the London South Bank University. Tool
wear, component forces, surface roughness and run-out were recorded and analysed. The
surface integrity study was mainly based on the physical and metallographic examination as
well as analysis of the microhardness variation of the machined surface and subsurface of the
workpiece materials using scanning electron microscope (SEM) and hardness measuring
equipment.
127
3.2 Work Material
The workpiece material used in this investigation is a commercially available alpha-beta
titanium-base, Ti-6Al-4V, alloy. The dimension of the Ti-6Al-4V workpiece bar is 300 mm
diameter x 300 mm long. The bars were previously homogenised and annealed. The nominal
chemical composition and physical properties of the workpiece material are given in Tables
3.1 and 3.2 respectively. Prior to the actual turning tests, the work material bars were trued,
centred and cleaned by removing up to 2 mm thickness of material at the top surface of the
workpiece in order to eliminate any surface defect that can adversely affect the machining
result. Each bar was mounted in the machine tool chuck and supported at the other end by the
tailstock to minimise vibration during the machining trials.
Table 3.1 - Nominal chemical composition of Ti-6Al-4V alloy (wt. %)
Chemical composition (wt. %)
Al V Fe O C H N Y Ti
Min. 5.50 3.50 0.30 0.14 0.08 0.01 0.03 50 ppm Balance
Max. 6.75 4.50 0.23
Table 3.2 - Physical properties of Ti-6Al-4V alloy.
Tensile
strength
(MPa)
0.2%
Proof
stress
(MPa)
Elongation
(%)
Density
(g cm
P
-3
P
)
Melting
point
(°C)
Measured
hardness
(C.I. – 99%)*
HV100
Thermal
conductivity at
20ºC
(W m
P
-1
P
KP
-1
P
)
900-1160 830 8 4.50 1650 Min.= 341
Max. = 363
6.6
* CI: Confidence interval of 99 %, represented by the minimum (Min.) and maximum (Max.) values.
3.3 Machine Tool
All the machining trials were carried out on a Colchester Electronic MASTIFF lathe
(Figure 3.1) equipped with a CNC control driven by an 11 kW motor driver which provides a
torque of 1411 Nm. The lathe has a variable spindle speed from 16-1800 rpm and a variable
feed rate ranging from 0.01 to 5.0 mm rev
-1
.
128
Figure 3.1 - Colchester Electronic MASTIFF CNC lathe.
3.4 Cutting Fluid
The cutting fluid used in the machining trials for both conventional and high pressure
systems is the HOCUT 3380, manufactured by Houghton, with a concentration of 6%. It is a
high lubricity emulsion coolant containing alkanolamine salts of fatty acids and
dicyclohexylamine. Additionally, this coolant contains anti-foaming and non splitting
properties, making it ideal for high pressure delivery applications. The required concentration
of coolant for all the machining trials was prepared by diluting the concentrate with water.
The coolant concentration was monitored with a pocket refractometer prior to machining and
at various stages of the machining trials. Cutting fluid was applied by flooding or
conventional coolant flow (CCF) to the cutting interface (overhead cooling) via a pipe with
nozzle diameter of 6 mm at an average flow rate of 2.7 l min
-1
(conventional method - D
129
direction (Figure 2.42)) and at a pressure lower than 0.3 MPa (3 bar) with CNC lathe coolant
delivery pump system. The distance from the nozzle to the cutting zone was kept at 100 mm.
3.5 High Pressure Unit
The high pressure coolant-jet was delivered by a high powered pumping unit detached
from the machine lathe. The high pressure pumping coolant system, Chipblaster (CV26-
3000), can operates at maximum load capacity of 22.3 kW, with a tank capacity of 454 l and
maximum flow rate of 197 l min
-1
and with a maximum output pressure of 21 MPa (210 bar).
The Chipblaster has four basic components: a reservoir (tank), a filtration system, an internal
high pressure pump and a control unit (Figure 3.2). A dual-filtration system filters the cutting
fluid before it enters the unit´s internal reservoir unit. The cutting fluid is pumped from the
high pressure pump to a special tool holder (Figure 3.3) which delivers the jet to the region
where the chip breaks contact in the tool (tool-chip interface - B direction in Figure 2.42). The
tool holder has an external nozzle with diameter of 2 mm and is kept at a distance of 13.5 mm
from the cutting zone. Three different high jet-pressures with the flow rates were used in the
trials:
• High pressure coolant (HPC) flow of 7 MPa (70 bar) with a flow rate of 16.9 l min
-1
(Figure 3.3);
• High pressure coolant flow of 11 MPa (110 bar) with a flow rate of 18.5 l min
-1
.
• High pressure coolant flow of 20.3 (203 bar) MPa with a flow rate of 24 l min
-1
.
Figure 3.2 - The high pressure pumping coolant system - Chipblaster (CV26-3000).
130
Figure 3.3 - Special tool holder and a cutting fluid jet-pressure of 7 MPa supply.
3.6 Argon Delivery System
The argon gas delivery system consists of an argon cylinder, a valve and a hose which
was connected to the same tool holder used in the high pressure cutting fluid delivery system.
The cylinder was placed in the back of the machine lathe as shown in Figure 3.4 (a). The
argon gas was delivered to the machining environment (tool-chip interface - B direction in
Figure 2.42) through a hose connected to a valve (Figure 3.4 (b)) where the flow rate of the
gas could be adjusted and supplied to the cutting interface through a nozzle at a constant flow
rate of 12 l min
-1
. Thermal conductivity of argon at a temperature of 300K (27°C) is 0.0177
Wm
-1
K
-1
.
Special
tool holder
131
(a) (b)
Figure 3.4 - Argon gas delivery system: (a) cylinder and (b) close-up view of the valve and
the hose.
3.7 Tool Material and Machining Procedure
All the cutting tool materials (Figure 3.5) used when machining Ti-6Al-4V alloy were
produced by Seco tools, with the exception of the micron-grain size ceramic (provided by
NTK tools manufacturer) and nano-grain size ceramics (currently under development). The
inserts have 2 different geometries: 80º rhomboid shape and square shape. Rhomboid shape
inserts were used when machining Ti-6Al-4V alloy with high pressure cutting fluid supplies.
These include: uncoated cemented carbide coded as (T1) 883 and (T2) 890 grades, coated
cemented carbide (T3) CP 200 and (T4) CP 250 grades, PCD (grade 20 with two different
grain sizes coded as PCD-STD (T5) and PCD-MM (T6), CBN (CBN 10 (T7), PCBN 300
(T8) and PCBN 300-P TiAlN/TiN coated (T9) grades), silicon carbide (SiC
w
) whisker
reinforced alumina ceramic (WG300) (T10). These insert grades are among the best Seco
grades for machining titanium alloys. Only the cemented carbide tools have an integral
132
obstruction-type chip-breaker on the rake faces. PCD inserts have a nose radius of 1.2 mm
whereas all other cutting tools have a nose radius of 0.8 mm. The square shape inserts were
used when machining Ti-6Al-4V alloy under conventional coolant supply only. They include:
micron-grain size silicon carbide (SiC
w
) whisker reinforced alumina ceramic (WG300) (T11)
and nano-grain size ceramic inserts AlB
2
BOB
3
B(mixed alumina) (T12) and SiB
3
BNB
4
B (silicon nitride
based) (T13).
Figure 3.5 - Cutting tools used in the machining trials: uncoated carbides: T1 (883 grade), T2
(890 grade), coated carbides T3 ( CP 200), T4 (CP 250 grade); PCD: T5 (20 grade with grain
size of 10 µm), T6 (20 grade with grain size < 10 µm); CBN: T7 (10 grade), PCBN: T8 (300
grade), PCBN: T9 (300-P grade); silicon carbide (SiC
w
) whisker reinforced alumina ceramic
inserts (WG300): T10 (rhomboid shaped) and T11 (square shaped); nano-grain ceramic
inserts: T12 (Al
2
O
3
grade) and T13 (Si
3
N
4
grade).
Table 3.3 gives the chemical and mechanical properties of the all rhomboid shape insert
grades and of T11 (SiC
w
whisker reinforced alumina ceramic inserts) square shaped cutting
tool materials used for the machining trials. The mechanical properties and nominal chemical
composition of the nano-grain ceramic inserts are given in Table 3.4. Prior to machining, an
insert was checked for any physical damage at the cutting edge with a travelling microscope
at a magnification of x20.
(T1) (T2) (T3) (T4)
(T5) (T6
)
(T7) (T8) (T9)
(T11) (T12) (T13)
(T10)
133
Table 3.3 - Specification, chemical and mechanical properties of the cutting tool materials
used in the machining trials.
Cemented Carbide (K grade)
Tool
material
Uncoated Coated
PCD
Insert
CBN
Solid PCBN
Whisker reinforced
ceramic
Tool code T1 T2 T3 T4 T5 T6 T7 T8 T9 T10 T11
Grade 883 890 CP200 CP250 STD MM 10 300 300-P WG300
ISO
designation
CNMG
120412-
M1
CNMG
120412-
MR3
CNMG
120412-
M1
CCMW
120408F-
L1
CNGA
120408S-
LO
CNMN
120412S
CNGN
120412TN
WA1
(rhomboid
shaped)
SNMG
120412
(square
shaped)
Chemical
composition
(wt.%)
93.8%
WC
6% Co
0.2%
(Ta,
Nb)C
93.7% WC
6% Co
0.3 Cr
2
C
3
93.8%
WC
6% Co
0.2%
(Ta, Nb)C
Diamond
+ Co
residue
50%
CBN
50% TiC
ceramic
(vol.)
90% CBN
10% Al
ceramic (vol.)
70%
(Al
2
O
3
, Y
2
O
3
, Zr
2
O)
30% SiC
Composition
and
thickness of
coating
(wt.%)
- -
TiAlN (3.5 µm)
TiN (0.5 µm
- - -
TiAlN
(3.5µm)
TiN
(0.5µm)
-
Hardness
(Knoops)
GPa
13
(1760
HV)
1753
(HV)
13
(1760
HV)
50.0 27.5 31.5
94.5 HRA
(2000
HVB
5
B)
Density
(g cm
-3
)
14.95 14.92 14.95 4.12 4.31 4.28 3.42 3.7
Thermal
conductivity
at 20ºC
(W m
-1
K
-1
)
110 - - 540 459 44 138 32
Substrate
grain size
(µm)
1.0 0.68 1.0 10 <10 2 - -
Whiskers
with
0.5µm of
diameter
Table 3.4 - Mechanical properties and chemical composition (wt%) of nano-ceramic tools
material (square shape inserts).
Tool
code Grade
ISO
designation
Hardness
(HVB
5
B)
Edge
toughness
(MPa mP
1/2
P
)
AlB
2
BOB
3
B SiC SiB
3
BNB
4
B TiCN YB
2
BOB
3
B ZrOB
2
B
T12 SAZT2 1779 10.54 75.0 - - 20.0 - 5.0
T13 SNCTN1
SNMG
120412
1670 6.92 4.5 4.5 68.3 18.2 4.5 -
134
Special tool holders were designed according to ISO designation for each cutting tool
material (rhomboid shape inserts) when machining with high pressure coolant (HPC) supply
Figures 3.6 (a), (b) and (c). These will enable the cutting fluid to travel from the high pressure
unit through its body and to be delivered at high pressure to the region where the chip breaks
contact with the tool via a specially designed orifice in its nozzle cap. The specially designed
tool holders were also used when machining with conventional coolant flow (CCF). A
different holder was employed when machining with square shape inserts (Figure 3.6 (d)). All
tool holders have a 25 mm shank. The effective geometry of the rhomboid inserts after rigid
clamping in the tool post provide: approach angle of 95°, back rake angle of -6°, side rake
angle of -6°, side relief angle of 6° and clearance angle of 6°, whereas the effective geometry
of the square inserts is an approach angle of 40°, back rake angle of -5°, side rake angle of 0°
and clearance angle of 6°.
Figure 3.6 - Tool holders used in the machining trials: (a) designation PCLNR2525-M12 used
for carbide tools (T1,T2,T3,T4); (b) designation SCLCR2525-M12 used for PCD tools
(T5,T6); (c) designation DCLNR2525-M12 used for CBN/PCBN and ceramic tools (T7,T8,
T9 and T10); (d) designation MSLNR-252512 used for square tools: micron-grain and nano-
grain size ceramics (T11,T12,T13).
(a) (b) (c) (d)
135
3.8 Cutting Conditions
Machining conditions for finish turning of Ti-6Al-4V alloy in this study started with
those usually employed in the aerospace industry. A 15 minutes tool life was chosen as the
benchmark for establishing acceptable cutting conditions during the turning trials. The initial
trials with uncoated cemented carbide (grade 883) tools were aimed at establishing high speed
conditions that will consistently achieve 15 minutes tool life under conventional coolant
supply, based on the stipulated tool wear rejection criterion: average flank wear,
VB
B
≥ 0.3 mm. The highest speed conditions achieved were used as the baseline conditions
for machining with various grades of advanced tools such as PCD, CBN/PCBN and nano-
grain ceramic tool materials. A constant feed rate and depth of cut of 0.15 mm rev
-1
and 0.5
mm respectively were used in all the machining trials. All the experimental trials and their
sequence are summarised in Table 3.5. This table shows the tools, machining environments,
cutting speed range and output variables investigated in this study. Tools T1, T2, T3, T4, T5,
T6, T7, T8, T9 and T10 have rhomboid-shape geometry and were used with conventional and
high pressure coolant supplies. The pressure of 7 MPa was employed with uncoated carbide
(T1) and PCD (T5) tools in order to investigate the performance of medium pressure in the
machinability of Ti-6Al-4V alloy. Argon gas supply was used with uncoated carbide (T1) and
coated carbide (T2) tools. Micron-grain size (T11) and nano-grain size (T12, T13) ceramic
tools have a square-shaped geometry and were used only with conventional coolant supply.
The output variables shown in Table 3.5 and listed below were monitored during the
machining trials, except the run-out which was measured at end of tool life when further
machining was halted:
i) Tool life (min);
ii) Tool wear (mm);
iii) Component forces (cutting and feed forces (N));
iv) Surface Roughness (R
a
- micron);
v) Runout (mm);
vi) Chip form.
Other output variables shown in Table 3.5 such as worn tool micrographs,
microhardness measurements, machined surfaces and subsurface examinations were evaluated
after each completed machining trial and after sample preparation.
136
Up to 450 experimental trials were carried out in this study, and it is important to note
that replications of experimental trials were performed for some machining conditions. For
such cases, average of tool lives was recorded. Additionally, tests were always repeated on
occurrence of abnormal and unpredictable tool failure.
Table 3.5 - Summary of the experimental tests carried out when finish turning of Ti-6Al-4V
alloy at a constant feed rate and depth of cut of 0.15 mm rev
-1
and a of 0.5 mm respectively.
Machining
Environment
Out-put variables
Stage Tool
CCF
HPC 7 MPa
HPC 11 MPa
HPC 20.3 MPa
Argon
Cutting speed
(m min
-1
)
Tool life
Tool wear
Component
forces
Worn tools
micrographs
Surface
roughness
Run-out
Microhardness
Machined surface
anal
y
sis
Subsurface
analysis
Chip photographs
Benchmark
trials
T1 X
60,90, 100,
110,120,130
X
T1 X X X X X
100,110,120,
130
X X X X X X X X X X
T2,T3 X X X 110, 130 X X X X X X X X X
T4 X X X X
100,110,120,
130
X X X X X X X X X X
T5 X X X X
140**,
150**,
160**,175,
200,230,250*
X X X X X X X X X X
T6 X X X
140**,150**,
160**, 175,
200*,230*,
250*
X X X X X X X X X X
T7,T8,
T9
X X X
150,200*,
250*
X X X X X X X X X X
T10 X X X X X
140,200,400,
500
X X X X X X X X X X
Actual
machining
trials
T11,
T12,
T13
X 110,130,200 X X X X X X X
Keys:
- CCF: Conventional coolant flow delivery system;
- HPC: High pressure coolant delivery system;
- *: Speed identified was not tested under conventional coolant supply;
- **: Speed identified was not tested under high pressure coolant supply(ies);
137
3.9 Tool Life Criteria
The tool life rejection criteria for finish turning operation employed in this investigation
are listed below:
i) Average flank wear, VB
B
≥ 0.3 mm;
ii) Maximum flank wear, VB
Bmax
≥ 0.4 mm;
iii) Nose wear, VC ≥ 0.3 mm;
iv) Notching at the depth of cut line, VN ≥ 0.6 mm;
v) Excessive chipping (flaking) or catastrophic fracture of the cutting edge;
vi) Surface roughness value (Ra) ≥ 1.6 µm (Centre line average);
vii) Run-out: 0.1 mm (100 µm)
3.10 Tool Wear Measurement
Tool wear measurement (flank wear, notch wear and nose wear) were carried out at
various intervals using a Mitutoyo tool maker´s microscope (Figure 3.7) connected to a digital
micrometer XY table with resolution of 0.001 mm at a magnification of x20. The insert is
removed from the tool holder and mounted on a small tool makers hand vice placed on the
microscope prior to measurement.
138
Figure 3.7 - Mitutoyo tool maker´s microscope.
3.11 Component Force Measurement
The component forces (cutting force, F
c
and feed force, F
f
) generated during the cutting
trials were recorded at the start of machining (after one minute machining time when the
cutting edge has not undergone pronounced wear) with the aid of a three component Kistler
piezoelectric tool post dynamometer. The electric signals generated by the component forces
during machining were fed into a charge amplifier connected to the dynamometer (Figure
3.8). This amplifier magnifies the electric signals (voltage) which can be read on a digital
oscilloscope. The maximum, minimum, peak-to-peak and average voltage values of the
cutting and feed forces were read directly from the screen. These values are converted to
mechanical units (Newton) multiplying them by with the charge amplification factor.
139
(a)
(b)
Figure 3.8 - (a) Kistler dynamometer for capturing forces generated during machining and
(b) Oscilloscope with charge amplifier.
3.12 Surface Roughness Measurement
Surface roughness values were recorded after one minute machining time using a
Surtronic-10 portable stylus type instrument that traverse over the machined surface (Figure
3.9 (a)). The relative displacement is highly magnified electronically and the results presented
as a surface roughness value (Ra) measured in micro metres (µm). The surface roughness
value was measured by positioning the instrument perpendicular to the feed marks on the
machined surface. The average of three readings at different locations on the workpiece bar
represents the surface roughness value of the machined surface. To ensure accuracy of the
readings the instrument was calibrated using a standard calibration block prior to use.
140
(a)
(b)
Figure 3.9: (a) Surtronic-10 portable stylus type used for surface roughness measurement;
(b) dial indicator Shockproof – BATY used for run-out measurement.
141
3.13 Runout Measurement
Run-out measurements are taken at the end of cut using a dial indicator Shockproof –
BATY which was placed in the cross slide of the toolholder and pointed towards the
workpiece bar as shown in Figure 3.9 (b). A small tension was applied and then the clock was
set in the zero position. Finally, the workpiece was rotated to determinate the run-out
variation.
3.14 Tool and Workpiece Specimen Preparation
The worn inserts are cleaned in acetone to remove oil stains, dust and adhering work
material on the surfaces to be examined. They are mounted on aluminium stubs with sticking
carbon glue, prior to examination in the scanning electron microscope (SEM), Hitachi S530,
as illustrated in Figure 3.10 (a). The worn surfaces can be viewed in either a three
dimensional or two dimensional view from any chosen angle. Relevant images of the worn
tools were selected and captured on a computer connected to the SEM. Analysis of these
micrographs helps to identify the failure mode(s) of the insert and to formulate possible wear
mechanism(s) involved. Additional information on the wear mechanism analysis can be
obtained by sectioning the worn region of the cutting tool with a diamond sliting saw which
will ensure minimum heat affected zone on the sectioned sample.
For specimens of the machined surfaces, 3 mm thickness was taken from the bars at the
end of tool life under various cutting conditions by grooving to a depth of 3 mm and then
cutting horizontally with a hand hacksaw. Each sample was further divided into two, one for
surface texture examination, and the other for subsurface alterations analysis after fine
polishing to 0.25 µm and etching in 3 ml of nitric acid, 6 ml of hydrochloric acid (HCl) and
100 ml of distilled water for 30 seconds in order to reveal the microstructure. The
microstructure is examined and photographed with a Nikon metallurgical optical microscope
(OPTIHOT-100) with attached camera (Figure 3.10 (b)). Specimens used to examine the
surface texture were first cleaned with acetone to remove oil stains and dirt sticking on the
surface prior to examination for any physical damage on the machined surfaces in a scanning
electron (SEM) microscope at high magnifications.
142
(a)
(b)
Figure 3.10 - (a) Hitachi (S530) Scanning Electron Microscope; (b) Nicon Metallurgical
Optical Microscope (OPTIPHOT-100) with computerised image system.
All samples used for examination of surface and subsurface alterations were mounted in
a mould using conductive phenolic powder (Resin-4 HQ). A Bueler – Automatic Moulding
143
Machine (Figure 3.11 (a)) was used to mount specimens of the work material and the
machined surfaces. An Automatic Grinding/Polishing Equipment – Metaserv 2000 – was used
to grind and polish specimens of the workpiece at a speed of 100 rpm (Figure 3.11 (b)). The
mounted specimens were polished using a series of silicon carbide papers of decreasing
grit/grain sizes (240, 400, 600, 800 and 1200) using water as lubricant. The samples were
cleaned in acetone and further polished to 6, 3, 1.0 and 0.25 µm finish on flat plate polishers
covered with “NP NAP” cloth impregnated with alumina paste and distilled water as a
lubricant. The samples were then examined using optical and electron microscopes at various
magnifications.
(a)
(b)
Figure 3.11 - (a) Buehler Automatic Mounting Press (Simplimet 2000); (b) Automatic
Grinding/Polishing Equipment (Metaserv 2000).
3.15 Microhardness Measurements below the Machined Surface
Microhardness of the workpiece samples sectioned after the machining trials were
measured using a micro-hardness testing machine, Mitutoyo (MVK – VL), at a magnification
of x55 and with an applied load of 0.1 Kgf after polishing the samples to 6µm finish (Figure
3.12). The microhardness of the workpiece sample taken prior to the start of the machining
trials is measured at various locations close to the central portion of the sample. This ensures
that any effect on the hardness of the workpiece material due to the initial cleaning and
sectioning of the sample is kept to a minimum/eliminated. The recorded values were treated
144
statistically and a confidence interval (C.I.) of 99 % (Table 3.2), represented by the minimum
(Min.) and maximum (Max.) Vickers hardness values recorded for the Ti-6Al-4V alloy bars
used for the machining trials. The microhardness measurements of the workpiece samples
were measured starting from a depth of 0.05 mm below the machined surface up to 2.0 mm.
Figure 3.12 -Mitutoyo (MVK – VL) Vickers micro-hardness tester machine.
CHAPTER IV
EXPERIMENTAL RESULTS
4.1 Benchmark trials - Machining of Ti-6Al-4V alloy with Uncoated Carbide (883 grade)
inserts
Figure 4.1 shows the average flank wear values recorded when machining with
uncoated cemented carbide (grade 883) tools at various cutting speeds under conventional
coolant supply, for 15 minutes cutting time. Average flank wear was the dominant tool failure
mode when machining Ti-6Al-4V alloy with this insert grade at the conditions investigated in
the benchmark trials. Average flank wear, VB
B
= 0.3 mm, was, then, the tool life rejection
criterion for the benchmark trials. It can be seen that the average flank wear rate increased
steadily with increase in cutting speed. Over 1800 % increase in wear was observed by
doubling the cutting speed from 60 to 120 m min
-1
. It can also be observed that at a speed of
120 m min
-1
, tool wear value is beyond the established rejection criterion of VB
B
= 0.3 mm.
Therefore, the cutting speed of 100 m min
-1
was chosen as the baseline cutting speed for the
actual machining trials.
146
0.0
0.2
0.4
0.6
0.8
1.0
Tool wear (mm)
60 90 100 110 120 130
Cutting speed (m/min)
Conventional (CCF)
Figure 4.1 - Average flank wear of uncoated carbide (T1) insert at various cutting speeds
under conventional coolant supply after 15 minutes machining time (benchmark trials).
4.2 Machining of Ti-6Al-4V alloy with various carbide tool grades (uncoated and coated
tools) under various machining environments
4.2.1 Tool Life
Figure 4.2 shows tool life (nose wear, VC ≥ 0.3 mm) recorded when machining Ti-6Al-4V
alloy with various grades of cemented carbide (T1, T2, T3, T4) inserts at various cutting
speeds and under various machining environments (conventional coolant flow, high pressure
coolant supplies of 7 MPa, 11 MPa and 20.3 MPa, and in an argon enriched environment).
Tool life of all the insert grades decreased with increasing cutting speed in all machining
environments investigated, as expected, due to a reduction in tool-chip and tool-workpiece
contact lengths and the consequent increase in both normal and shear stresses at the tool tip
(GORCZYCA, 1987). Reduction in tool-chip and tool-workpiece contact areas may also
concentrate the high temperature generated to a relatively smaller area as well as shifting the
highest temperature region closer to the cutting edge. This phenomenon combines with higher
stresses acting at the cutting edge to promote softening of the cutting tool and consequently
accelerates tool wear processes, thereby reducing tool life. Tool life generally increased with
increasing coolant pressure, especially when machining with the T1 and T4 tool grades.
Re
j
ection crite
r
ion: VB
B
= 0.3 m
m
147
0.0
10.0
20.0
30.0
40.0
50.0
60.0
70.0
80.0
100 110 120 130
Cutting speed (m/min)
Tool life (min) '
T1 (Argon) T1 (CCF) T1 (7 MPa) T1 (11 MPa) T1 (20.3 MPa)
T2 (CCF) T2 (11MPa) T2 (20.3MPa) T3 (CCF) T3 (11MPa)
T3 (20.3MPa) T4 (CCF) T4 (Argon) T4 (11 MPa) T4 (20.3 MPa)
Figure 4.2 - Tool life (nose wear, VC ≥ 0.3 mm) recorded when machining Ti-6Al-4V alloy
with different cemented carbide insert grades under conventional coolant flow (CCF), high
coolant pressures of 7 MPa, 11 MPa and 20.3 MPa and in argon enriched environment at
various speed conditions.
Machining in an argon enriched environment gave the worst performance in terms of
tool life relative to conventional and high pressure coolant supplies. It can also be seen in
Figure 4.2 that tools T1 (uncoated tool) and T4 (coated tool) with the same substrate
composition (Table 3.3) gave the best performance in terms of tool life in all conditions
investigated relative to tools T2 (uncoated tool) and T3 (coated tool) with the same substrate
composition. In general T4 tool gave the best performance of all the cemented carbide insert
grades tested. However, T1 tool outperformed T4 tool when machining under higher coolant
pressure of 20.3 MPa at a speed of 110 m min
-1
and in argon enriched environment. T2 and
T3 tools generally exhibited similar performance in terms of tool life at the cutting conditions
investigated, except when machining with the highest coolant pressure of 20.3 MPa at a speed
of 110 m min
-1
, where over 65% increase in tool life was achieved with T3 tool. The longest
tool life of 73 min was recorded when machining at the lowest speed of 100 m min
-1
with T1
tool under a coolant pressure of 7 MPa. The second highest tool life was also recorded with
T1 tool using the highest pressure supply of 20.3 MPa at a cutting speed of 110 m min
-1
. T4
T1
T4 T1
T2 T3
T4
T1
T4
T1
T2
T3
T4
148
tool gave the second best performance, in terms of tool life, relative to T2 and T3 tools when
machining in all the machining environments investigated at a lower cutting speed of
110 m min
-1
. T2 and T3 inserts gave similar performance when machining at a higher cutting
speed of 130 m min
-1
. Figure 4.2 also shows that increasing coolant pressure up to 11 MPa
improved tool life for all the carbide tool grades at the highest speed of 130 m min
-1
.
Machining with 20.3 MPa coolant pressure, however, produced slightly lower tool life
relative to those obtained with 11MPa coolant pressure at a speed of 130 m min
-1
, Table 4.1.
The highest gains in tool life were achieved when machining with T4 tool at speeds of
110 m min
-1
and 120 m min
-1
with 20.3 MPa coolant pressure where 356% and 291%
improvement in tool lives were recorded, respectively. Up to 84% and 246% improvement in
tool lives were achieved when machining with T1 and T4 tools, respectively, under most
aggressive conditions of 130 m min
-1
using 20.3 MPa coolant pressure. Over 165%
improvement in tool life was achieved when machining with T1 tool with 7 MPa coolant
pressure relative to conventional coolant supply at the lowest cutting speed of 100 m min
-1
whereas, at cutting speeds in excess of 100 m min
-1
, improvement in tool life remained
constant (at about 35%). These results show that tool life increased with increasing coolant
pressure for all the insert grades investigated. The ranking order for carbide tools in terms of
average gain in tool life relative to conventional coolant flow is T4, T3, T2 and T1 (Table
4.1). This also shows that the coolant pressure has a significant effect on tool wear pattern and
hence recorded tool life when machining Ti-6Al-4V alloy with carbide tools under finishing
conditions.
Figure 4.2 and Table 4.1 also show significant reduction in tool life when machining in
an argon enriched environment relative to conventional coolant supply. Average reduction in
tool life for T1 and T4 tools are 47 % and 44% respectively. It can also be seen that the
uncoated carbide T1 tool outperformed the coated carbide T4 tool when machining with
conventional coolant flow and in an argon enriched environment at the cutting conditions
investigated. A marginal improvement in tool life of 8% was achieved when machining with
T1 tool grade relative to T4 in all the environments tested at the lowest speed of 100 m min
-1
.
Machining with T1 tool at a cutting speed of 110 m min
-1
produced up to 65% and 74%
improvements in tool life with an argon enriched environment and with conventional coolant
flow, respectively, compared to T4 grade.
149
Table 4.1 - Percentage improvement in tool life relative to conventional coolant supply after
machining Ti-6Al-4V alloy with different grades of carbides.
Coolant pressure (MPa) Tool
Speed
(m min
-1
)
Argon
7 11 20.3
100 -50 165.7 . .
110 -48.3 31.3 63.2 195
120 -51.5 36.1 78.5 86.1
T1
130 -40.2 35.4 130.5 84.1
110 . . 102.4 133.7
T2
130 . . 138 128
110 . . 76.1 250
T3
130 . . 205.2 176.3
100 -50,2 . . .
110 -50.9 . 256.4 356.4
120 -45.6 . 272.1 291.1
T4
130 -28 . 272 246
4.2.2 Tool wear when machining Ti-6Al-4V alloy with various carbide insert grades
Flank wear and nose wear are typical wear patterns observed when machining Ti-6Al-
4V alloy with carbide tools. The flank wear rate is generally lower than the nose wear rates.
Figure 4.3 shows plots of nose wear rates of different cemented carbide insert grades when
machining Ti-6Al-4V alloy under various cutting environments. Machining with carbide tools
under finishing conditions showed steady increase in nose wear rate with increasing cutting
speed. It can also be seen that nose wear rate generally decreases with increasing coolant
pressure. The lowest nose wear rates were recorded when machining with T1 and T4 tools
with 11 MPa coolant pressure. Rapid increase in nose wear rate was recorded when
machining with T1 and T4 tools at the higher speed conditions of 130 m min
-1
under
conventional coolant flow and in the presence of argon. The highest nose wear rate was
recorded when machining Ti-6Al-4V alloy with T4 tool grade in the presence of argon in all
cutting conditions investigated. This gave increased nose wear rate of about 38%, 462% and
310% relative to conventional coolant flow, coolant pressures of 11 MPa and 20.3 MPa,
150
respectively when machining at 130 m min
-1
. High nose wear rate was also recorded when
machining with T1 tool grade in presence of argon at a speed of 130 m min
-1
. This increase is
about 74%, 144%, 479% and 252% relative to conventional coolant flow and coolant
pressures of 7 MPa, 11 MPa and 20.3 MPa, respectively. Nose wear rate increased with
prolong machining in all the cutting environments tested when machining with the carbide
insert grades. At the initial stages of cutting, nose wear was generally uniform for inserts used
for machining with high coolant pressures. Further machining accelerated the nose wear as
illustrated in Figures 4.4 and 4.5. Figure 4.4 shows a reduction in nose wear with increasing
coolant pressure for both T1 and T4 insert grades. The lowest nose wear was recorded when
machining with T1 tool with the highest coolant pressure of 20.3 MPa while machining with
T4 tool in argon enriched environment gave the highest nose wear. Figure 4.5 shows that the
highest and lowest nose wear was recorded with T2 and T3 insert grades at a speed of
110 m min
-1
using conventional and 20.3 MPa coolant pressure, respectively.
0.00
0.02
0.04
0.06
0.08
0.10
0.12
100 110 120 130
Cutting speed (m/min)
Nose wear rate (mm/min)
T1 (CCF) T1 (11MPa) T1 (20.3 MPa)
T2 (CCF) T2 (11MPa) T2 (20.3 MPa)
T3 (CCF) T3 (11MPa) T3 (20.3 MPa)
T4 (CCF) T4 (11MPa) T4 (20.3 MPa)
T1 (7 MPa) T1 (Argon) T4 (Argon)
Figure 4.3 - Nose wear rate curves of different cemented carbide insert grades when
machining Ti-6Al-4V alloy under conventional coolant flow (CCF), high coolant pressures of
7 MPa, 11 MPa and 20.3 MPa and in argon enriched environment, at a feed rate of
0.15 mm rev
-1
and a depth of cut of 0.5 mm.
151
0,0
0,1
0,2
0,3
0,4
0,5
0 102030405060
Time (min)
Nose wear (mm)
T1 (CCF) T4 (CCF)
T1 (Argon) T4 (Argon)
T1 (7 MPa) T4 (11 MPa)
T1 (11 MPa) T4 (20.3 MPa)
T1 (20.3 MPa)
Figure 4.4 - Nose wear curves when finish machining with cemented carbide (T1 and T4)
inserts at a cutting speed of 110 m min
-1.
0.0
0.1
0.2
0.3
0.4
0.5
0 102030405060
Time (min)
Nose wear (mm
)
T2 (CCF) T3 (CCF)
T2 (11 MPa) T3 (11 MPa)
T2 (20.3 MPa) T3 (20.3 MPa)
Figure 4.5 - Nose wear curves when finish machining with cemented carbide (T2 and T3)
inserts at a cutting speed of 110 m min
-1.
152
Figures 4.6 – 4.10 show worn cutting edges obtained after machining Ti-6Al-4V alloy
with T1 tool grade (uncoated carbide tool – 883 grade) under various machining conditions.
The T1 grade experienced regular nose wear when machining under conventional coolant
flow at the lower speed of 100 m min
-1
(Figure 4.6 (a)) and more severe nose wear at a speed
of 130 m min
-1
(Figure 4.6 (b)). Figures 4.6 (a) and (b) also illustrate the extent of adhesion of
work material to the nose of the worn tool. Figures 4.7 to 4.9 are the micrographs of T1 worn
insert after machining with 7 MPa, 11 MPa and 20.3 MPa coolant pressures at cutting speeds
of 110 m min
-1
, 120 m min
-1
and 130 m min
-1
, respectively. There is the evidence of adhesion
of work material to the tool nose as result of the intermittent contact between the tool and the
work material during machining. Flank wear rate when machining under high coolant
pressures was generally very low relative to nose wear rate. Figure 4.10 shows a smoothly
worn tool with slight cracks on a worn T1 insert after machining in the presence of argon at a
speed of 130 m min
-1
.
(a) Tool life = 27.4 min (b) Tool life = 8.2 min
Figure 4.6 - Worn T1 insert after machining Ti-6Al-4V alloy with conventional coolant
supply at a speed of (a) 100 m min
-1
and (b) 130 m min
-1
.
Rake face
Nose wear
Rake face
Flank wear
Nose wear
153
Tool life = 23.9 min
Figure 4.7 - Wear generated at the cutting edge of uncoated carbide T1 insert after machining
Ti-6Al-4V alloy with a coolant pressure of 7MPa at a speed of 110 m min
-1
.
Tool life = 23.2 min
Figure 4.8 -Wear generated at the cutting edge of uncoated carbide T1 insert after machining
Ti-6Al-4V alloy with a coolant pressure of 11MPa at a speed of 120 m min
-1
.
Nose wear
Rake face
Adhesion
Nose wear
Rake face
Adhesion
154
Tool life = 15.1 min
Figure 4.9 - Worn cutting edge of uncoated carbide T1 insert after machining Ti-6Al-4V alloy
under a coolant pressure of 20.3 MPa at a speed of 130 m min
-1
.
Tool life = 4.9 min
Figure 4.10 - Worn cutting edge of uncoated carbide T1 insert after machining Ti-6Al-4V
alloy in argon enriched environment at a speed of 130 m min
-1
.
Figures 4.11-4.13 are typical micrographs of worn T2 (uncoated carbide tool – 890
grade) inserts after machining Ti-6Al-4V alloy at different cooling and cutting speed
conditions. Both flank and nose wears occurred simultaneously at the cutting edge. The flank
Adhesion
Rake face
Nose wear
Rake face
Nose wear
Flank wear
155
wear tend to be displaced towards the nose region with prolonged machining, making only the
nose wear responsible for tool rejection. Irregular flank and nose wear patterns were observed
across the worn flank face when machining with conventional flow and at the highest coolant
pressure of 20.3 MPa at a cutting speed of 130 m min
-1
(Figure 4.13).
Tool life = 5 min
Figure 4.11 - Flank and nose wears at the cutting edge of uncoated carbide T2 insert grade
after machining Ti-6Al-4V alloy under conventional coolant supply at a speed of 130 m min
-1
,
a feed rate of 0.15 mm rev
-1
and a depth of cut of 0.5 mm.
(a) Tool life = 16.8 min (b) Tool life = 11.9 min
Figure 4.12 -Wear generated at the cutting edge of uncoated carbide T2 insert after machining
Ti-6Al-4V alloy with a coolant pressure of 11 MPa at a speed of (a) 110 m min
-1
and (b) 130
m min
-1
.
Rake face
Nose wear
Flank wear
Nose wear
Rake face
Rake face
Nose wear
156
Tool life = 11.4 min
Figure 4.13 - Wear generated at the cutting edge of uncoated carbide T2 insert after
machining Ti-6Al-4V alloy with a coolant pressure of 20.3 MPa at a speed of 130 m min
-1
.
Figures 4.14-4.16 are micrographs of T3 worn tools (coated carbide tool – CP200
grade) after machining Ti-6Al-4V alloy with different coolant pressures and cutting speeds.
These figures show that T3 tool grade experienced more severe flank and nose wears in all
machining environments tested. Visible grooves were observed on the worn inserts as clearly
illustrated in Figures 4.15 and 4.16 (a) and (b).
(a) Tool life = 9.2 min (b) Tool life = 3.8 min
Figure 4.14 - Worn cutting edge of T3 coated carbide insert when machining with
conventional coolant supply at a speed of (a) 110 m min
-1
and (b) 130 m min
-1
.
Rake face
Rake face
Nose wear
Nose wear
Rake face
157
Tool life = 16.2 min
Figure 4.15 - Flank and nose wears a the cutting edge of T3 coated carbide insert after
machining Ti-6Al-4V alloy with a coolant pressure of 11 MPa at a speed of 110 m min
-1
.
(a) Tool life = 32.2 min (b) Tool life = 10.5 min
Figure 4.16 - Flank and nose wears at the cutting edge of T3 coated carbide insert after
machining Ti-6Al-4V alloy with a coolant pressure of 20.3 MPa at a speed of (a) 110 m min
-1
and (b) 130 m min
-1
.
Figures 4.17-4.20 are micrographs of T4 worn tools (coated carbide tool – CP250
grade) after machining Ti-6Al-4V alloy with different cooling conditions. Figure 4.17 shows
the flank wear spreading along the flank face and displaced towards the nose region. Figure
4.18 is an enlarged view of the worn tool after machining with a coolant pressure of 11 MPa
at a cutting speed of 110 m min
-1
illustrating the extent of adhesion of work material to the
tool nose. This was also observed in Figures 4.17 and 4.19 suggesting that high temperatures
may be generated at the cutting interfaces during machining. Figure 4.20 (a) shows a
Nose wear
Flank wear
Rake face
Rake face
Rake face
Nose wear
158
smoothly worn tool obtained after machining in the presence of argon at a speed of
100 m min
-1
while Figure 4.20 (b) shows that the T4 insert experienced pronounced flank
wear and crater wears when machining in an argon enriched environment at a speed of
120 m min
-1
. These figures also show that the crater wear developed very close to the cutting
edge.
Tool life = 5 min
Figure 4.17 - Worn cutting edge of T4 coated carbide insert after machining Ti-6Al-4V alloy
with conventional coolant supply at a speed of 130 m min
-1
.
Tool life = 39.2 min
Figure 4.18 - Adhesion of work material on a worn T4 coated carbide insert after machining
Ti-6Al-4V alloy with a coolant pressure of 11 MPa at a speed of 110 m min
-1
.
Rake face
Nose wear
Adhesion
Nose wear
Adhesion
159
(a) Tool life = 50.2 min (b) Tool life = 30.9 min
Figure 4.19 - Nose wear at the cutting edge of T4 coated carbide insert after machining Ti-
6Al-4V alloy with a coolant pressure of 20.3 MPa at a speed of (a) 110 m min
-1
and
(b) 120 m min
-1
.
(a) Tool life = 12.7 min (b) Tool life = 4.3 min
Figure 4.20 - Wear at the cutting edge of T4 coated carbide insert after machining Ti-6Al-4V
alloy in argon enriched environment at a speed of (a) 100 m min
-1
and (b) 120 m min
-1
.
4.2.3 Component forces when machining with various carbide insert grades
Figures 4.21 and 4.22 show variations in cutting and feed forces, respectively, when
machining Ti-6Al-4V alloy with different grades of carbide tools under various cutting speeds
and machining environments. The component forces were recorded at the beginning of cut
when the cutting edge has not undergone pronounced wear. Figure 4.21 suggests that the
cutting forces generally decreased with increase in coolant pressure when machining with T1
Rake face Rake face
Nose wear
Nose wear
Adhesion
Rake face
Rake face
Nose wear
Nose wear
Flank face
Crater wear
160
and T4 insert grades and generally increase with increase in coolant pressure when machining
with T2 and T3 insert grades. Figure 4.21 also shows that in general, cutting forces marginally
increased with increasing speed when machining with T1 and T2 in all cutting environments
tested, unlike T3 and T4 insert grades. Cutting forces recorded with T2 tool grade were
generally higher than those recorded with T3 tool in all conditions tested. High cutting forces
were generated when machining with T1 and T4 tool grades in argon environment relative to
conventional coolant flow.
0
90
180
270
360
100 110 120 130
Cutting speed (m/min)
Cutting force, Fc (N) '
T1 (CCF) T1 (11MPa) T1 (20.3 MPa)
T2 (CCF) T2 (11MPa) T2 (20.3 MPa)
T3 (CCF) T3 (11MPa) T3 (20.3 MPa)
T4 (CCF) T4 (11MPa) T4 (20.3 MPa)
T1 (7 MPa) T1 (Argon) T4 (Argon)
Figure 4.21 - Cutting forces (Fc) recorded at the beginning of cut when machining Ti-6Al-4V
alloy with different cemented carbide grades under various cutting conditions.
Recorded feed forces generally increase marginally with an increase in cutting speed
(Figures 4.22). This figure also shows that feed forces increased with an increase in coolant
pressure when machining with T1, T2 and T3 tool grades, unlike T4 tool. It can be seen from
Figure 4.22 that the highest feed forces were generated when machining with T1 tool grade
under a coolant pressure of 7 MPa. Machining with T4 tool grade in the presence of argon
provided the lowest feed forces in all conditions tested.
161
0
90
180
270
360
100 110 120 130
Cutting speed (m/min)
Feed force, Ff (N)
T1 (CCF) T1 (11MPa) T1 (20.3 MPa)
T2 (CCF) T2 (11MPa) T2 (20.3 MPa)
T3 (CCF) T3 (11MPa) T3 (20.3 MPa)
T4 (CCF) T4 (11MPa) T4 (20.3 MPa)
T1 (7 MPa) T1 (Argon) T4 (Argon)
Figure 4.22 - Feed forces (F
f
) recorded at the beginning of cut when machining Ti-6Al-4V
alloy with different cemented carbide grades under various cutting conditions.
4.2.4 Surfaces roughness and runout deviation when machining with various carbide insert
grades
Figure 4.23 and 4.24 show the surface roughness and circular runout deviation values,
respectively, recorded when machining Ti-6Al-4V alloy with different grades of carbides at
various cutting speeds and under various cutting environments. It can be seen from Figure
4.23 that the surface roughness values recorded in all the conditions investigated varied
between 0.3 and 1.0 µm, well below the stipulated rejection criterion of 1.6 µm. This figure,
however, shows evidence of deterioration of the surface finish when machining at higher
speed conditions.
162
0.0
0.3
0.6
0.9
1.2
100 110 120 130
Cutting speed (m/min)
Surface roughnes
s
Ra ( m)
T1 (CCF) T1 (11MPa) T1 (20.3 MPa)
T2 (CCF) T2 (11MPa) T2 (20.3 MPa)
T3 (CCF) T3 (11MPa) T3 (20.3 MPa)
T4 (CCF) T4 (11MPa) T4 (20.3 MPa)
T1 (7 MPa) T1 (Argon) T4 (Argon)
Figure 4.23 - Surface roughness values recorded at the beginning of cut when machining Ti-
6Al-4V alloy with different cemented carbide grades under various cutting conditions.
Runout tolerances are used to control the functional relationship of one feature to
another or a feature to a datum axis. This tolerance is applicable to rotating parts where the
composite surface criterion is based on the part function and design requirements. When
dealing with three-dimensional objects, circular runout is defined as the amount that is
allowed to deviate from the central axis at one cross section (POLLACK, 1988). Figure 4.24
shows that runout values recorded in all the conditions investigated varies between 2 and 14
µm. These values are far lower than the stipulated rejection criterion value of 100 µm.
Machining with T4 tool grade provided lower runout values in all conditions investigated
compared to other tool grades. Runout values recorded increased with increasing cutting
speed when machining with T2 and T3 insert grades under conventional coolant flow while
no variation was observed when machining with T2 and T3 inserts under high coolant
pressures.
163
0
3
6
9
12
15
100 110 120 130
Cutting speed (m/min)
Runout (
m)
T1 (CCF) T1 (11MPa) T1 (20.3 MPa)
T2 (CCF) T2 (11MPa) T2 (20.3 MPa)
T3 (CCF) T3 (11MPa) T3 (20.3 MPa)
T4 (CCF) T4 (11MPa) T4 (20.3 MPa)
T1 (7 MPa) T1 (Argon) T4 (Argon)
Figure 4.24 – Runout variation recorded at the end of cut when machining Ti-6Al-4V alloy
with different cemented carbide grades under various cutting conditions.
4.2.5 Surfaces generated after machining with various carbide insert grades
Figures 4.25-4.29 show micrographs of surfaces generated when machining with
uncoated carbide T1 tool grade under various machining environments and at different cutting
speeds. Surfaces generated consist of well-defined and uniform feed marks running
perpendicular to the direction of relative work-tool motion with no evidence of plastic flow.
There is evidence of localised incipient melting of the machined surfaces when machining at a
higher cutting speed of 130 m min
-1
under conventional coolant flow (Figure 4.25 (b)).
(a)
(b)
Figure 4.25 - Surfaces generated after machining with uncoated carbide T1 tool with
conventional coolant supply at cutting speeds of (a) 110 m min
-1
and (b) 130 m min
-1
.
164
(a) (b)
Figure 4.26 - Surfaces generated after machining with uncoated carbide T1 tool with a coolant
pressure of 7 MPa at cutting speeds of (a) 100 m min
-1
and (b) 130 m min
-1
.
(a) (b)
Figure 4.27 - Surfaces generated after machining with uncoated carbide T1 tool with a coolant
pressure of 11 MPa at cutting speeds of (a) 110 m min
-1
and (b) 120 m min
-1
.
(a) (b)
Figure 4.28 - Surfaces generated after machining with uncoated carbide T1 tool with a coolant
pressure of 20.3 MPa at cutting speeds of (a) 120 m min
-1
and (b) 130 m min
-1
.
165
(a) (b)
Figure 4.29 - Surfaces generated after machining with uncoated carbide T1 tool in an argon
enriched environment at cutting speeds of (a) 110 m min
-1
and (b) 120 m min
-1
.
Figures 4.30 and 4.31 show micrographs of surfaces generated when machining with
uncoated carbide (T2) and PVD coated carbide (T3) tool grades under high pressure coolant
supplies of 11MPa and 20.3 MPa at a cutting speed of 110 m min
-1
. There are well-defined
uniform feed marks running perpendicular to the direction of relative work-tool motion.
(a) (b)
Figure 4.30 - Surfaces generated after machining with uncoated carbide T2 tool with coolant
pressures of (a) 11 MPa and (b) 20.3 MPa at a cutting speed of (a) 110 m min
-1
.
166
(a) (b)
Figure 4.31 - Surfaces generated after machining with coated carbide T3 tool with coolant
pressures of (a) 11 MPa and (b) 20.3 MPa at a cutting speed of 110 m min
-1
.
Typical surfaces generated when machining with PVD coated carbide (T4) tool grade
under various machining environments at a cutting speed of 120 m min
-1
as shown in Figure
4.32. Well-defined uniform feed marks running perpendicular to the direction of relative
work-tool motion with no evidence of plastic flow can be seen. No surface tears and chatter
marks were observed after machining Ti-6Al-4V alloy with various cemented carbide tool
grades. Generally machining with all the carbide grades under high pressure coolant supplies
generated acceptable machined surfaces, conforming to the standard specification established
for machined aerospace components (Rolls-Royce CME 5043).
167
(a) (b)
(c) (d)
Figure 4.32 - Surfaces generated after machining with coated carbide T4 tool with
(a) conventional coolant supply, (b) in argon enriched environment, (c) coolant pressure of
11 MPa and (d) 20.3 MPa at a cutting speed of 120 m min
-1
.
4.2.6 Surface hardness after machining with various carbide tool grades
Figures 4.33-4.42 are plots of the variations of microhardness values recorded from the
top of the machined surface up to about 1.5 mm below the machined surface. Note that the
range of measured values on the graphs is demarcated by confidence interval (C.I.),
represented by the minimum (Min.) and maximum (Max.) Vickers hardness values recorded
for Ti-6Al-4V alloy bars prior to machining. This is because the hardness of alloys varies
within a range of values. The range (C.I.) is used here instead of the average hardness because
it gives a realistic assessment of the material hardness.
168
Figures 4.33-4.37 are plots of microhardness values of machined surfaces after
machining with uncoated carbide T1 tool grade under various machining environments. The
plots show relatively low variation in hardness when machining with T1 tool grade with
conventional coolant flow at cutting speeds up to 110 m min
-1
(Figure 4.33). There is,
however, evidence of surface hardening beyond the bulk hardness of material at the top
surface when machining at speeds in excess of 110 m min
-1
. Surface hardening up to about
0.4 mm below the machined surfaces of Ti-6Al-4V alloy was observed after machining with
uncoated carbide T1 tool grade at a cutting speed of 120 m min
-1
. Softening of the machined
surfaces also occurred, especially, at the higher speed of 130 m min
-1
. Figure 4.34 shows
evidence of softening at the top surface up to about 0.15 below the machined surfaces after
machining with T1 tool with a coolant pressure of 7 MPa at speeds of 100 m min
-1
and
120 m min
-1
. This figure also shows evidence of surface hardening up to about 0.3 mm below
the machined surfaces when machining at a cutting speed of 110 m min
-1
. Figure 4.35 shows
the evidence of softening of machined surfaces after machining with T1 tool with a coolant
pressure of 11 MPa at speeds up to 120 m min
-1
. However, evidence of hardening of the
machined surfaces up to about 0.4 mm depth was observed when machining at highest cutting
speed of 130 m min
-1
. Machining with T1 tool grade under the highest coolant pressure of
20.3 MPa gave minimum hardness variation with a uniform distribution of hardness values
within the confidence interval of hardness values prior to machining (Figure 4.36). This
figure, generally, shows softening of the machined surfaces when machining at all the cutting
speeds investigated. Figure 4.37 shows microhardness values of machined surfaces after
machining with T1 tool grade in the presence of argon. It can be seen that hardness values are
generally distributed within the confidence interval of the hardness values.
169
300
340
380
420
460
0,00 0,20 0,40 0,60 0,80 1,00 1,20 1,40 1,60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 100 m/min V = 110 m/min
V = 120 m/min S = 130 m/min
Min._Value (341 HV) Max._Value (363 HV)
Figure 4.33 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide (T1)
insert grade with conventional coolant supply.
300
340
380
420
460
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 100 m/min V = 110 m/min
V = 120 m/min S = 130 m/min
Min_Value (341 HV) Max_Value (363 HV)
Figure 4.34 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide (T1)
insert grade with 7 MPa coolant pressure.
170
300
340
380
420
460
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 110 m/min Min._Value (341 HV)
V = 120 m/min Max._Value (363 HV)
V = 130 m/min
Figure 4.35 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide (T1)
insert grade with 11MPa coolant pressure.
240
280
320
360
400
440
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 110 m/min Min._Value (341 HV)
V = 120 m/min Max._Value (363 HV)
V = 130 m/min
Figure 4.36 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide (T1)
insert grade under 20.3 MPa coolant pressure.
171
300
330
360
390
420
0.00 0.30 0.60 0.90 1.20 1.50 1.80 2.10
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 100 m/min V = 110 m/min
V = 120 m/min V = 130 m/min
Min_Value (341 HV) Max_Value (363 HV)
Figure 4.37 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide (T1)
insert grade in argon-enriched environment.
Figure 4.38 is a plot of the variation of microhardness values recorded at various
locations below the machined surfaces after machining with uncoated carbide T2 tool grade
using conventional coolant flow and high pressure coolant supplies of 11 MPa and 20.3 MPa.
The recorded hardness values are generally distributed within the confidence interval of
hardness values of the work material prior to machining. However, Figure 4.38 also shows
evidence of surface hardening up to about 0.25 mm below the machined surface when
machining with conventional coolant supply at the lowest speed of 100 m min
-1
. It can also be
seen in Figure 4.38 that surface hardness decreased with increasing cutting speed up to about
0.25 mm depth when machining with conventional coolant supply.
Figure 4.39 is a plot of the variation of microhardness values recorded at various
locations below the machined surfaces after machining with coated carbide T3 tool grade with
conventional coolant flow and high pressure coolant supplies of 11 MPa and 20.3 MPa. It can
be seen that the hardness values are uniformly distributed within the confidence interval of the
hardness values prior to machining.
172
300
340
380
420
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 100 m/min (CCF) V = 100 m/min (HP 20.3MPa)
V = 130 m/min (CCF) V = 130 m/min (HP 20.3MPa)
V = 100 m/min (HP 11MPa) Min_Value (341 HV)
V = 130 m/min (HP 11MPa) Max_Value (363 HV)
Figure 4.38 - Hardness variation after machining Ti-6Al-4V alloy with uncoated carbide (T2)
insert grade under various cutting conditions.
300
340
380
420
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 100 m/min (CCF) V = 100 m/min (HP 20.3MPa)
V = 130 m/min (CCF) V = 130 m/min (HP 20.3MPa)
V = 100 m/min (HP 11MPa) Min_Value (341 HV)
V = 130 m/min (HP 11MPa) Max_Value (363 HV)
Figure 4.39 - Hardness variation after machining Ti-6Al-4V alloy with coated carbide (T3)
insert grade under various cutting conditions.
173
Figures 4.40-4.42 are plots of microhardness values of machined surfaces after
machining with coated carbide T4 tool grade under various machining environments. These
figures show that machining with T4 tool grade in all machining environments investigated
reduced microhardness of the machined surfaces. This suggests the softening of machined
surface up to about 1 mm below the machined surface when machining with conventional
coolant flow at all cutting speeds investigated, since all the hardness values recorded are
below the minimum hardness values of bulk hardness of the material (Figure 4.40). The least
softening depth (235 HV) was recorded when machining at the lowest cutting speed of
100 m min
-1
. Softening of machined surfaces was also observed when machining under high
pressure coolant supplies (Figures 4.41 and 4.42). This softening was more pronounced when
machining at the highest cutting speed of 130 m min
-1
.
210
250
290
330
370
0.00 0.20 0.40 0.60 0.80 1.00 1.20
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 100 m/min Min._Value (341 HV)
V = 110 m/min Max._Value (363 HV)
V = 120 m/min
Figure 4.40 - Hardness variation after machining Ti-6Al-4V alloy with coated carbide (T4)
insert grade under conventional coolant supply.
174
270
300
330
360
390
420
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 110 m/min Min._Value (341 HV)
V = 120 m/min Max._Value (363 HV)
V = 130 m/min
Figure 4.41 - Hardness variation after machining Ti-6Al-4V alloy with coated carbide (T4)
insert grade under 11MPa coolant pressure supply.
210
250
290
330
370
410
450
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 110 m/min Min._Value (341 HV)
V = 120 m/min Max._Value (363 HV)
V = 130 m/min
Figure 4.42 - Hardness variation after machining Ti-6Al-4V alloy with coated carbide (T4)
insert grade under 20.3 MPa coolant pressure supply.
175
4.2.7 Subsurface micrographs after machining Ti-6Al-4V alloy with various carbides insert
grades
A scanning electron microscope was employed to view etched cross-sections of the
machined surfaces in order to detect microstructural alteration and other damages on the
machined surfaces. Figures 4.43-4.57 are microstructures of etched machined surfaces of Ti-
6Al-4V alloy after machining with different grades of uncoated (T1 and T2) and coated
carbide (T3 and T4) inserts under various cutting conditions. All the micrographs exhibit
similar characteristics. The well defined grain boundaries are clear evidence that there was no
microstructure alteration such as plastic deformation, in the subsurface of machined surfaces.
(a) Tool life = 18.2 min (b) Tool life = 8.2 min
Figure 4.43 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide T1
insert under conventional coolant supply at cutting speeds of (a) 110 m min
-1
and
(b) 130 m min
-1
.
(a) Tool life = 72.8 min (b) Tool life = 17.7 min
Figure 4.44 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide T1
insert under a coolant pressure of 7 MPa at cutting speeds of (a) 100 m min
-1
and
(b) 120 m min
-1
.
176
(a) Tool life = 29.7 min (b) Tool life = 23.2 min
Figure 4.45 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide T1
insert under a coolant pressure of 11 MPa at cutting speeds of (a) 110 m min
-1
and
(b) 120 m min
-1
.
(a) Tool life = 24.2 min (b) Tool life = 15.1 min
Figure 4.46 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide T1
insert under a coolant pressure of 20.3 MPa at cutting speeds of (a) 120 m min
-1
and
(b) 130 m min
-1
.
(a) Tool life = 9.4 min (b) Tool life = 6.3 min
Figure 4.47 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide T1
inserts in an argon enriched environment at cutting speeds of (a) 110 m min
-1
and
(b) 120 m min
-1
.
177
(a) Tool life = 8.3 min (b) Tool life = 5.0 min
Figure 4.48 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide T2
inserts with conventional coolant supply at a cutting speed of (a) 110 m min
-1
and
(b) 130 m min
-1
.
(a) Tool life = 16.8 min (b) Tool life = 11.9 min
Figure 4.49 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide T2
inserts with a coolant pressure of 11 MPa at cutting speeds of (a) 110 m min
-1
and
(b) 130 m min
-1
.
(a) Tool life = 19.4 min (b) Tool life = 11.4 min
Figure 4.50 - Microstructure of Ti-6Al-4V alloy after machining with uncoated carbide T2
tools with a coolant pressure of 20.3 MPa at cutting speeds of (a) 110 m min
-1
and
(b) 130 m min
-1
.
178
(a) Tool life = 9.2 min (b) Tool life = 3.8 min
Figure 4.51 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T3
inserts with conventional coolant supply at a cutting speed of (a) 110 m min
-1
and
(b) 130 m min
-1
.
(a) Tool life = 16.2 min (b) Tool life = 11.6 min
Figure 4.52 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T3 tools
with a coolant pressure of 11 MPa at cutting speeds of (a) 110 m min
-1
and (b)
130 m min
-1
.
(a) Tool life = 19.4 min (b) Tool life = 11.4 min
Figure 4.53 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T3 tools
with a coolant pressure of 20.3 MPa at cutting speeds of (a) 110 m min
-1
and (b) 130 m min
-1
.
179
(a) Tool life = 25.5 min (b) Tool life = 7.9 min
Figure 4.54 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T4 tools
with conventional coolant supply at cutting speeds of (a) 100 m min
-1
and
(b) 120 m min
-1
.
(a) Tool life = 29.4 min (b) Tool life = 18.6 min
Figure 4.55 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T4 tools
with a coolant pressure of 11 MPa at cutting speeds of (a) 120 m min
-1
and (b) 130 m min
-1
.
(a) Tool life = 50.2 min (b) Tool life = 17.3 min
Figure 4.56 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T4 tools
with a coolant pressure of 20.3 MPa at cutting speeds of (a) 110 m min
-1
and (b) 130 m min
-1
.
180
(a) Tool life = 4.3 min (b) Tool life = 3.6 min
Figure 4.57 - Microstructure of Ti-6Al-4V alloy after machining with coated carbide T4 tools
in an argon enriched environment at cutting speeds of (a) 120 m min
-1
and
(b) 130 m min
-1
.
4.2.8 Chips shapes
Figure 4.58 contains samples of chips generated when machining with different grades
of uncoated and coated carbide tools (T1,T2,T3 and T4) under various machining conditions.
Generally chips produced with all the tools with conventional coolant flow are snarled shapes
(Figures 4.58 (a), (b), (f), (h) and (j)). Machining with T1 insert with conventional coolant
flow produced long continuous tubular type chips, unlike the small segmented chips produced
when machining under high pressure coolant supplies of 11 MPa and 20.3 MPa (Figures 4.58
(d) and (e)). Machining with T1 insert with a coolant pressure of 7 MPa and with T2 insert
with high pressure coolant of 11 MPa produced either segmented or snarled type chips. The
nature of the chips produced with higher pressure coolant supplies is markedly different from
that obtained with conventional coolant flow, as illustrated in Figure 4.58 (c) and (g).
Segmented type chips were produced when machining with T4 tool grade under high coolant
pressure supplies, similar to those produced with T1 tool grade.
181
(a) T1 (conventional)
V = 110 m min
-1
(b) T1 (argon)
V = 110 m min
-1
(c) T1 (7 MPa)
V = 110 m min
-1
(d) T1 (11 MPa)
V = 110 m min
-1
(e) T1 (20.3 MPa)
V = 120 m min
-1
(f) T2 (conventional)
V = 110 m min
-1
(g) T2 (HP 11MPa)
V = 110 m min
-1
(h) T3 (conventional)
V = 110 m min
-1
(i)
T3 (20.3 MPa)
V = 110 m min
-1
11 MPa 20.3 MPa
(j) T4 (conventional)
V = 110 m min
-1
(k) T4 (argon)
V = 110 m min
-1
(l) T4
V = 110 m min
-1
Figure 4.58 - Chips generated when machining Ti-6Al-4V alloy with different carbide tool
grades under various cutting conditions: (a) continuous tubular chip; (b), (f), (h) and (k)
continuous and snarled chips, (c), (g) and (i) partially segmented chips, (d), (e) and (l)
segmented C-shaped chips.
182
4.3 Machining of Ti-6Al-4V alloy with different grades of PCD tools under various
coolant supply pressures
4.3.1. Tool Life
Figure 4.59 shows tool life (nose wear, VC ≥ 0.3 mm) recorded when machining Ti-
6Al-4V alloy with the standard (T5) and the multi-modal (T6) grades of PCD inserts at
various cutting speeds and at various coolant supply pressures investigated. Tool life of all the
PCD insert grades decreased with increase cutting speed in all machining environments
investigated due to a reduction in tool-chip and tool-workpiece contact length and the
consequent increase in both normal and shear stresses at the tool tip (GORCZYCA, 1987).
Lower tool life was recorded when machining with both grades of PCD inserts under
conventional coolant flow. Tool life generally increased with increasing in coolant pressure
when machining with the larger grain size (T5) grade, unlike lower tool life recorded when
machining with the smaller grain size (T6) grade at the highest pressure of 20.3 MPa relative
to 11 MPa. Note that high pressure coolant supplies were employed at cutting speeds in
excess of 160 m min
-1
. However, machining with T5 PCD grade with 7MPa coolant pressure
gave longer tool life than with conventional coolant flow in all the cutting speeds investigated
and with 11 MPa coolant pressure at cutting speeds of 175 and 230 m min
-1
. This is illustrated
in Table 4.2, which provides a summary of the percentage improvement in tool life when
machining with T5 and T6 insert grades under various coolant pressures relative to
conventional coolant supply. Encouraging tool life of 115 and 97 minutes was recorded when
machining with the larger grain size (T5) tool grade at a cutting speed of 175 m min
-1
under
coolant supply pressures of 20.3 MPa and 7 MPa, respectively. These represent over 20 and
17 folds improvement in tool life, respectively, relative to conventional coolant flow. The
highest gain in tool life was achieved when machining under most aggressive conditions of
230 m min
-1
with 7 MPa and 20.3 MPa coolant supplies when over 20 fold improvement in
tool life was recorded (Table 4.2). Additionally, improvement in tool life using T5 insert with
11 MPa coolant supply increased with increase in cutting speed.
183
0.0
20.0
40.0
60.0
80.0
100.0
120.0
140 150 160 175 200 230 250
Cutting speed (m/min)
Tool life (min) '
T5 (CCF) T5 (7MPa) T5 (11MPa) T5 (20.3MPa)
T6 (CCF) T6 (11MPa) T6 (20.3MPa)
Figure 4.59 - Tool life (nose wear, VC ≥ 0.3 mm) recorded when machining Ti-6Al-4V alloy
with PCD-STD (T5) and PCD MM (T6) tool grades with conventional coolant flow (CCF)
and high coolant pressures of 7 MPa, 11 MPa and 20.3 MPa at various cutting speed
conditions.
Table 4.2 - Percentage improvement in tool life relative to conventional coolant supply after
machining Ti-6Al-4V alloy with PCD inserts (STD and MM grades).
Coolant Pressure (MPa)
Tool
Speed
(m min
-1
)
7 11 20.3
175 1701.8 1081.5 2038.9
200 1096.7 1146.7 1583.3
T5
230 2230 1650 2040
175 . 879.2 574.5
T6
200 . 1956.5 1313
It can be also seen from Figure 4.59 that smaller grain size (T6) tool grade exhibited
superior performance, in terms of tool life, than T5 tool grade when machining with
184
conventional coolant flow and at a high pressure coolant supply of 11 MPa at cutting speeds
up to 175 m min
-1
and 250 m min
-1
, respectively. Over 74% improvement in tool life was
achieved with T6 tool grade at a speed of 150 m min
-1
with conventional coolant flow relative
to T5 tool. Over 42% and 26% improvement in tool life were achieved when machining with
T6 tool grade at speeds of 175 and 200 m min
-1
under 11MPa coolant supply compared with
T5 insert. Higher tool lives were recorded when machining with T5 tool grade using the
highest pressure coolant of 20.3 MPa than with pressure of 11 MPa at cutting speeds up to
230 m min
-1
. Up to 20 and 13 fold improvement in tool life were recorded when machining
with T6 insert under 11 MPa and 20.3 MPa coolant supplies, respectively, relative to
conventional coolant flow at a speed of 200 m min
-1
(Figure 4.59). Machining at a lower
speed of 175 m min
-1
with T6 insert grade under 11 MPa and 20.3 MPa coolant supplies,
respectively, gave only about 9 and 6 fold improvement relative to conventional coolant flow.
From these results it is clear that coolant pressure has a significant effect on tool wear pattern
and hence recorded tool life when machining Ti-6Al-4V alloy with PCD tools under finishing
conditions. Additionally, the performance of PCD tools depends on the cutting speed
employed. Machining with PCD (T5) tool grade at a speed of 140 m min
-1
under conventional
coolant flow gave up to 6 folds improvement in tool life relative to uncoated carbide (T1)
insert after machining at a cutting speed of 130 m min
-1
(Figure 4.2) and over 10 fold
improvement when machining in an argon enriched environment.
4.3.2 Tool wear when machining Ti-6Al-4V alloy with different grades of PCD tools
Flank wear and nose wear are typical wear patterns observed when machining Ti-6Al-
4V alloy with PCD inserts. Flank wear rate is generally lower than the nose wear rate, hence
nose wear was the predominant failure mode in all the cutting conditions investigated. Figure
4.60 shows typical nose wear rate curves when machining Ti-6Al-4V alloy with T5 and T6
grades of PCD inserts under conventional coolant flow and various coolant pressure supplies.
There was a steady increase in nose wear with increasing cutting speed when machining with
PCD tools under finishing conditions. Figure 4.60 also shows that there was negligible
difference in nose wear rate when machining with both T5 and T6 insert grades at the coolant
pressures investigated at higher cutting speeds up to 250 m min
-1
.
185
0.00
0.05
0.10
0.15
0.20
0.25
0.30
0.35
140 150 160 175 200 230 250
Cutting speed (m/min)
Nose wear rate (mm/min)
T5 (CCF) T5 (7 MPa) T5 (11MPa)
T5 (20.3 MPa) T6 (CCF) T6 (11MPa)
T6 (20.3 MPa)
Figure 4.60 - Nose wear rate when machining Ti-6Al-4V alloy with PCD inserts with
conventional coolant flow and high coolant pressures of 7 MPa, 11 MPa and 20.3 MPa at
various cutting speed conditions.
Nose wear rates increased with prolong machining using both PCD insert grades under
all coolant supply pressures employed (Figure 4.61). At the initial stages of cutting, nose wear
were generally uniform when machining under high coolant pressures. Further machining
rapidly increased wear land. The curves in Figure 4.61 also shows that more severe nose wear
was recorded when machining with larger grain size (T5) insert grade under conventional
coolant flow unlike the very low, uniform and gradual nose wear rate recorded when
machining at the highest pressure coolant supply of 20.3 MPa. Machining with 11 MPa
coolant supply pressure, however, gave higher nose wear than with 7 MPa pressure (Figure
4.61). Uniform nose wear with prolong machining, up to 60 min, was observed when
machining under 11 MPa pressure. This is increased rapidly after 60 min cutting time until the
end of tool life (64 min). Figure 4.61 also shows that higher nose wear occurred when
machining with the T5 grade at cutting speed of 175 m min
-1
under conventional coolant flow
relative to the smaller grain size (T6) insert. T5 insert, however, outperformed T6 tool grade
in terms of lower nose wear with prolong machining at high coolant pressures.
186
0.0
0.1
0.2
0.3
0.4
0.5
0 15 30 45 60 75 90 105 120
Time (min)
Nose wear (mm
)
T5 (CCF) T6 (CCF)
T5 (7 MPa) T6 (11 MPa)
T5 (11 MPa) T6 (20.3 MPa)
T5 (20.3 MPa)
Figure 4.61 - Nose wear when finish machining with PCD-STD (T5) and PCD-MM inserts
grades (T6) at a cutting speed of 175 m m
-1.
Figures 4.62–4.68 show worn PCD tools after finish turning of Ti-6Al-4V alloy at
various machining conditions. Figures 4.62 (a) and (b) show the worn cutting edges of a T5
insert after machining at cutting speeds of 140 m min
-1
and 200 m min
-1
, respectively, under
conventional coolant supply. A uniform nose wear, spreading along the flank face can be
observed. Although there was the absence of a well-defined built-up-edge on the worn inserts,
the presence of an adherent interfacial layer (adhesion of the workpiece material) on the rake
and flank faces of worn T5 and T6 inserts when machining mainly with conventional coolant
flow was observed (Figures 4.62). Significant crater wear was observed on both grades of
PCD tools when machining with high pressure coolant supplies (Figures 4.63, 4.67 and 4.68).
The crater wear usually occurs very close to the cutting edge and joined the notched region at
the end of the depth of cut region when machining with T5 insert with 7MPa and 11 MPa
pressures at a cutting speed of 175 m min
-1
(Figures 4.63 (a) and 4.64 (a)). Adhesion of work
material can be seen on both the rake and flank faces of the worn T5 insert grade after
machining with 7MPa coolant pressure at higher speeds of 175 and 200 m min
-1
(Figure 4.63
(a) and (b)). Evidence of notching at the depth of cut region can also be seen on the worn T6
insert after machining at a speed of 175 m min
-1
with conventional coolant flow (Figure 4.66).
187
In addition to the pronounced crater wear, plucking of PCD tool particles was observed after
machining with higher coolant pressures. This process consists of small holes or pockets on
the worn tool surface due to mechanical removal of tiny particles from the surface. This
process tend to be accelerated at higher coolant pressures and is capable of removing any
loose tool material from the rake face with the potential to erode the brittle PCD tool material,
thereby exhibiting the characteristic rough texture as illustrated in Figures 4.63-4.68.
(a) Tool life = 56.7 min (b) Tool life = 3.0 min
Figure 4.62 - Wear observed on T5 insert after machining Ti-6Al-4V alloy with conventional
coolant supply at a speed of (a) 140 m min
-1
and (b) 200 m min
-1
.
(a) Tool life = 97.3 min (b) Tool life = 23.3 min
Figure 4.63 - Worn T5 insert after machining Ti-6Al-4V alloy with a 7MPa coolant pressure
and at a speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
Rake face
Rake face
Nose wear
Adhesion
Nose wear
Rake face
Nose wear
Adhesion
Crater wear
Nose wear
Adhesion
188
(a) Tool life = 63.8 min (b) Tool life = 17.5 min
Figure 4.64 - Worn T5 insert after machining Ti-6Al-4V alloy with 11 MPa coolant pressure
at a speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
(a) Tool life = 50.5 min (b) Tool life = 8.3 min
Figure 4.65 - Wear observed on a T5 insert after machining Ti-6Al-4V alloy with 20.3 MPa
coolant pressure at a speed of (a) 200 m min
-1
and (b) 250 m min
-1
.
Nose wear
Rake face
Adhesion
Crater wear
Nose wear
Rake face
Nose wear
Crater wear
Nose wear
Adhesion
189
(a) Tool life = 10.2 min
Figure 4.66 - Wear observed on a T6 insert after machining Ti-6Al-4V alloy with
conventional coolant supply at a speed of 175 m min
-1
.
(a) Tool life = 90.7 min (b) Tool life = 26.1 min
Figure 4.67 - Worn T6 insert after machining Ti-6Al-4V alloy with 11MPa coolant pressure at
a speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
Nose wear
Notching
Crater wear
Notching
190
(a) Tool life = 90.7 min (b) Tool life = 26.1 min
Figure 4.68 - Worn T6 insert after machining Ti-6Al-4V alloy with 20.3 MPa coolant
pressure at a speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
4.3.3 Component forces when machining with different grades of PCD tools
Figures 4.69 and 4.70 show cutting forces and feed forces, respectively, recorded at the
beginning of cut when machining Ti-6Al-4V alloy with PCD (T5 and T6) inserts with various
cutting speeds and coolant supply pressures. Figure 4.69 shows a norminal variation in cutting
forces when machining with PCD tools with all the coolant supply pressures investigated.
Cutting forces generally increased with increasing cutting speed when machining with T5 and
T6 inserts with conventional coolant supply. It can also be seen from Figure 4.69 that cutting
forces generally decreased with increasing coolant pressure, up to 11 MPa, when machining
with the larger grain size (T5) tool grade at higher cutting speeds up to 230 m min
-1
.
Machining with the smaller grain size (T6) insert generated lower cutting forces than
machining with T5 inserts. Marginal variation in cutting forces was observed when machining
with T6 insert with 11 MPa coolant pressure at the cutting speeds investigated. Increase in
cutting speed led to a slight reduction in cutting forces under 20.3 MPa coolant supply
pressure. Cutting forces generally decrease with increase in coolant pressure when machining
with T6 inserts.
Crater wear
Rake face
Nose wea
r
191
50
100
150
200
250
140 150 160 175 200 230 250
Cutting speed (m/min)
Cutting force, Fc (N)
T5 (CCF) T6 (CCF)
T5 (7 MPa) T6 (11MPa)
T5 (11MPa) T6 (20.3 MPa)
T5 (20.3 MPa)
Figure 4.69 - Cutting forces (Fc) recorded at the beginning of cut when machining Ti-6Al-4V
alloy with PCD-STD (T5) and PCD-MM insert grades (T6) at various cutting conditions.
50
100
150
200
250
140 150 160 175 200 230 250
Cutting speed (m/min)
Feed force, F
f
(N) '
T5 (CCF) T6 (CCF)
T5 (7 MPa) T6 (11MPa)
T5 (11MPa) T6 (20.3 MPa)
T5 (20.3 MPa)
Figure 4.70 - Feed forces (Ff) recorded at the beginning of cut when machining Ti-6Al-4V
alloy with PCD-STD (T5) and PCD-MM insert grades (T6) at various cutting conditions.
192
Feed forces generated when machining with T5 and T6 inserts were generally lower
than cutting forces especially at higher cutting speeds (Figure 4.70). The curves show that
feed forces generally decreased with increasing in cutting speed when machining with both
T5 and T6 inserts using conventional coolant flow and tend to increase at higher speeds when
machining under high coolant pressures. It is interesting to note that the highest feed forces
were recorded when machining with T6 insert at the highest coolant supply of 20.3 MPa.
4.3.4 Surfaces roughness and runout when machining with different grades of PCD tools
Figure 4.71 and 4.72 show the surface roughness and runout values, respectively,
recorded when machining Ti-6Al-4V alloy with PCD (T5 and T6) inserts at various cutting
conditions. Curves in Figure 4.71 show that surface roughness values recorded with PCD
tools in all the conditions investigated varies between 0.5 and 2.2 µm. Most of the recorded
values are, however, lower than the stipulated rejection criterion of 1.6 µm. Machining with
both T5 and T6 insert grades did not produce any significant changes in the surfaces finish
generated considering the anticipated scatter. The curves show gradual deterioration of the
surface finish with increase in coolant pressure when machining with T5 insert at cutting
speeds up to 230 m min
-1
while increase in coolant pressure improve surface finish generated
when machining with T6 insert grade. This figure also shows that good surface finish and
norminal variation in surface roughness values were obtained when machining with T5 insert
with conventional coolant flow. Machining with T6 insert with a coolant pressure of 11MPa
produced the highest surface roughness values at higher speeds in excess of 175 m min
-1
.
Figure 4.72 are plots of runout values recorded when machining Ti-6Al-4V alloy with
T5 and T6 inserts at various cutting conditions. The runout values vary between 2 and 12 µm.
These values are far lower than the stipulated rejection criterion value of 100 µm. Generally
runout values recorded decreases with increasing cutting speed when machining with both
grades of PCD tools. T5 tools gave lower runout values than T6 tools when machining with
all coolant supplies at cutting speeds up to 230 m min
-1
. The least runout value of 2 µm was
recorded when machining with T5 insert with high coolant pressure of 20.3 MPa at a cutting
speed of 200 m min
-1
193
0.0
0.5
1.0
1.5
2.0
2.5
140 150 160 175 200 230 250
Cutting speed (m/min)
Surface roughness,
Ra (
µ
m)
T5 (CCF) T6 (CCF)
T5 (7 MPa) T6 (11MPa)
T5 (11MPa) T6 (20.3 MPa)
T5 (20.3 MPa)
Figure 4.71- Surface roughness values recorded at the beginning of cut when machining Ti-
6Al-4V alloy with T5 and T6 inserts at various cutting conditions.
0
3
6
9
12
15
140 150 160 175 200 230 250
Cutting speed (m/min)
Runout (
m)
T5 (CCF) T6 (CCF)
T5 (7 MPa) T6 (11MPa)
T5 (11MPa) T6 (20.3 MPa)
T5 (20.3 MPa)
Figure 4.72 – Runout values recorded at the end of cut after machining Ti-6Al-4V alloy with
T5 and T6 insert grades at various cutting conditions.
194
4.3.5 Surface alteration after machining with different grades of PCD tools
Surfaces generated when machining the titanium alloy with PCD (T5 and T6) tools at
various cutting speeds and under various coolant supply pressure generally consist of well-
defined and uniform feed marks running perpendicular to the direction of relative work-tool
motion with no evidence of plastic flow (Figures 4.73-4.79).
Machining with the larger grain size T5 insert with a coolant pressure of 11 MPa did not
leave any damage on machined surfaces at all cutting speeds investigated, Figure 4.75. The
surface damages observed existed mainly as localised incipient melting of the machined
surfaces (Figures 4.73 (a), 4.74 (b), 4.76-4.79) and micro-pits especially when machining with
T6 insert using high coolant pressures (Figures 4.78 and 4.79). Most of machined surfaces
exhibit evidence of localised incipient melting of the machined surfaces when machining with
both grades of PCD tools at the different coolant supplies employed. Distribution of the
localised incipient melting of machined surfaces are worse when machining with T5 inserts
with the highest coolant pressure of 20.3 MPa (Figure 4.76) and more severe when machining
with T6 inserts with coolant supply pressures of 11 MPa and 20.3 MPa (Figures 4.78 and
4.79). The continuity of the feed marks is impaired by the localised incipient melting of the
machined surface when machining with T5 insert at a cutting speed of 175 m min
-1
using
conventional coolant flow (Figure 4.73 (a)), at a speed of 200 m min
-1
under a coolant
pressure of 7 MPa (Figure 4.74 (b)) and at the highest coolant pressure of 20.3 MPa. The
continuity of the feed marks on machined surfaces was also interrupted by the localised
incipient melting of machined surfaces when machining with T6 inserts at all the conditions
investigated (Figures 4.77-4.79). Irregular feed marks were observed in some cases when
machining with the smaller grain size T6 insert at a cutting speed of 175 m min
-1
with
conventional coolant supply and also with 11 MPa coolant pressure (Figures 4.77 (a) and 4.78
(a), respectively). Generally surfaces generated under all the finishing conditions investigated
when machining with both PCD tool grades were free from other damages such as cracking,
tearing and rupture that are detrimental to machined components. Therefore, surfaces
generated in these trials are acceptable and conform to the standard specification established
for machined aerospace components (Rolls-Royce CME 5043).
195
(a) (b)
Figure 4.73 - Surfaces generated after machining with PCD (T5) inserts with conventional
coolant supply at a cutting speed of (a) 175 m min
-1
and (b) 200 m min
-1
.
(a) (b)
Figure 4.74 - Surfaces generated after machining with PCD (T5) inserts with a coolant
pressure of 7 MPa at a cutting speed of (a) 175 m min
-1
and (b) 200 m min
-1
.
(a) (b)
Figure 4.75 - Surfaces generated after machining with PCD (T5) inserts with a coolant
pressure of 11 MPa at a cutting speed of (a) 200 m min
-1
and (b) 250 m min
-1
.
196
(a) (b)
Figure 4.76 - Surfaces generated after machining with PCD (T5) inserts with a coolant
pressure of 20.3 MPa at a cutting speed of (a) 175 m min
-1
and (b) 200 m min
-1
.
(a) (b)
Figure 4.77 - Surfaces generated after machining with PCD (T6) inserts with conventional
coolant supply at a cutting speed of (a) 175 m min
-1
and (b) 200 m min
-1
.
(a) (b)
Figure 4.78 - Surfaces generated after machining with PCD (T6) inserts with a coolant
pressure of 11 MPa at a cutting speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
197
(a) (b)
Figure 4.79 - Surfaces generated after machining with PCD (T6) inserts with a coolant
pressure of 20.3 MPa at a cutting speed of (a) 175 m min
-1
and (b) 200 m min
-1
.
4.3.6 Surface hardness after machining with different grades of PCD tools
Figures 4.80-4.86 are plots of the variations of microhardness values recorded from the
top up to about 1.5 mm below the machined surface after machining with different PCD tools
at various cutting speeds and under various coolant supply pressures. The plots indicate that
the hardness depth of the machined surface generally increased with increasing in cutting
speed when machining with the larger grain size (T5) tools using conventional coolant flow
and high coolant pressures supplies up to 11 MPa (Figures 4.80-4.83). A comparison of
Figures 4.80 and 4.84 shows that lower hardness values were recorded when machining
Ti-6Al-4V with T6 tools.
250
290
330
370
410
450
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 140 m/min V = 150 m/min
V = 160 m/min V = 175 m/min
V = 200 m/min V = 230 m/min
Min._Value (341 HV) Max._Value (363 HV)
Figure 4.80 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T5) insert with
conventional coolant supply.
198
250
290
330
370
410
450
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 175 m/min V = 200 m/min
V = 230 m/min V = 250 m/min
Min._Value (341 HV) Max._Value (363 HV)
Figure 4.81 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T5) insert with
7 MPa coolant pressure supply.
Analysis of Figures 4.82 and 4.85 suggests that there is considerable softening of the
surfaces generated after machining with a 11MPa coolant pressure using both insert grades
The use of a higher coolant pressure of 20.3 MPa resulted to pronounced softening when
machining with T5 inserts relative to T6 inserts (Figures 4.83 and 4.86).
250
290
330
370
410
450
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 175 m/min V = 200 m/min
V = 230 m/min V = 250 m/min
Min._Value (341 HV) Max._Value (363 HV)
Figure 4.82 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T5) insert with
11 MPa coolant pressure supply.
199
250
290
330
370
410
450
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 175 m/min V = 200 m/min
V = 230 m/min V = 250 m/min
Min._Value (341 HV) Max._Value (363 HV)
Figure 4.83 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T5) insert with
20.3 MPa coolant pressure supply.
200
240
280
320
360
400
440
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 140 m/min V = 200 m/min
V = 150 m/min Min._Value (341 HV)
V = 160 m/min Max._Value (363 HV)
V = 175 m/min
Figure 4.84 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T6) insert with
conventional coolant supply.
200
250
290
330
370
410
450
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 175 m/min V = 200 m/min
V = 230 m/min V = 250 m/min
Min._Value (341 HV) Max._Value (363 HV)
Figure 4.85 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T6) insert with
11 MPa coolant pressure supply.
200
250
300
350
400
450
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
V = 175 m/min V = 200 m/min
V = 230 m/min V = 250 m/min
Min._Value (341 HV) Max._Value (363 HV)
Figure 4.86 - Hardness variation after machining Ti-6Al-4V alloy with PCD (T6) insert with
20.3 MPa coolant pressure supply.
4.3.7 Subsurface alteration after machining Ti-6Al-4V alloy with different grades of PCD
tools
Figures 4.87-4.93 are microstructures of etched machined surfaces of Ti-6Al-4V alloy
after machining with different PCD (T5 and T6) inserts at the cutting conditions investigated.
201
These figures show no mechanical damage or microstructural alterations after machining with
all the carbide grades at the cutting conditions investigated.
(a) Tool life = 56.7 min (b) Tool life = 1 min
Figure 4.87 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T5) insert with
conventional coolant supply at a cutting speed of (a) 140 m min
-1
and (b) 230 m min
-1
.
(a) Tool life = 97.3 min (b) Tool life = 13.2 min
Figure 4.88 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T5) insert with a
coolant pressure of 7 MPa at a cutting speed of (a) 175 m min
-1
and (b) 250 m min
-1
.
202
(a) Tool life = 63.8 min (b) Tool life = 17.5 min
Figure 4.89 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T5) insert with a
coolant pressure of 11 MPa at a cutting speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
(a) Tool life = 50.5 min (b) Tool life = 21.4 min
Figure 4.90 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T5) insert with a
coolant pressure of 20.3 MPa at a cutting speed of (a) 200 m min
-1
and (b) 230 m min
-1
.
(a) Tool life = 69.9 min (b) Tool life = 2.3 min
Figure 4.91 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T6) insert with
conventional coolant supply at a cutting speed of (a) 140 m min
-1
and (b) 200 m min
-1
.
203
(a) Tool life = 90.7 min (b) Tool life = 26.1 min
Figure 4.92 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T6) insert with a
coolant pressure of 11 MPa at a cutting speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
(a) Tool life = 68.8 min (b) Tool life = 17.9 min
Figure 4.93 - Microstructure of Ti-6Al-4V alloy after machining with PCD (T6) insert with a
coolant pressure of 20.3 MPa at a cutting speed of (a) 175 m min
-1
and (b) 230 m min
-1
.
4.3.8 Chips shapes
Figures 4.94 (a)–(g) show different chips generated when machining with T5 and T6
grades of PCD tools under various cutting conditions. Machining Ti-6Al-4V alloy with the
larger grain size (T5) tool grade using conventional coolant flow produced long helical form
of chips (Figure 4.94 (a)), whereas T6 insert produced long continuous tubular type chips
(Figure 4.94 (e)). Machining with both T5 and T6 inserts with high pressure coolant supplies
produced smaller segmented chips (Figures 4.94 (b), (c), (d), (f) and (g)). It is important to
note that the nature of swarf produced when machining with high pressure coolant supplies is
markedly different to those obtained with conventional coolant flow. Coolant supply at high
204
pressures tend to enhance chip segmentation as the chip curl radius is reduced significantly,
hence maximum coolant pressure is restricted only to a smaller area on the chip.
(a) T5 (conventional)
V = 200 m min
-1
(b) T5 (7MPa)
V = 175 m min
-1
(c) T5 (11 MPa)
V = 200m min
-1
(d) T5 (20.3MPa)
V = 200 m min
-1
(e) T6 (Conventional)
V = 140 m min
-1
(f) T6 (11 MPa)
V = 175 m min
-1
(g) T5 (20.3MPa)
V = 200 m min
-1
Figure 4.94 - Chips generated when machining Ti-6Al-4V alloy with different grades of PCD
at various cutting conditions: (a): snarled chip; (e): long continuous chip, (b), (c), (d), (f) and
(g): segmented C-shaped chips.
4.4 Machining of Ti-6Al-4V alloy with different grades of CBN/PCBN tools under
various coolant supply pressures
4.4.1. Tool Life
Figure 4.95 shows tool life (VC ≥ 0.3 mm or VN ≥ 0.6 mm) recorded when machining
Ti-6Al-4V alloy with three different grades of CBN (T7) and PCBN (T8 and T9) inserts at
205
various cutting speeds and with conventional coolant flow and high pressure coolant supplies
of 11 MPa and 20.3 MPa. Tool life generally decreased with increasing cutting speed when
machining with the lower (50 vol.%) CBN content (T7 grade), unlike when machining with
PCBN (T8 and T9) tool grades with higher (90 vol.%) CBN content. Additionally, tool life of
the CBN grades generally increased with increasing the coolant pressure, especially when
machining with the T7 tool grade at speeds of 150 and 250 m min
-1
, which was more sensitive
to coolant pressure than others. T7 insert gave the best performance, in terms of tool life, at all
the conditions investigated. Over 68% and 150% improvement in tool life was recorded when
machining with tool T7 grade at a cutting speed of 150 m min
-1
with the coolant pressure
supplies of 11 MPa and 20.3 MPa, respectively, in comparison to conventional coolant flow.
However, tool life decreased with increasing the coolant pressure when machining at a cutting
speed in excess of 200 m min
-1
. In general T8 and T9 inserts exhibited similar performance in
terms of tool life (generally below 1 min) at the conditions investigated. Figure 4.95 also
shows that coating on the T9 insert did not influence tool performance.
0
1
2
3
4
5
150 200 250
Cutting speed (m/min)
Tool life (min) '
T7 (CCF) T7 (11MPa) T7 (20.3MPa)
T8 (CCF) T8 (11MPa) T8 (20.3MPa)
T9 (CCF) T9 (11MPa) T9 (20.3MPa)
T7 T8 T9
T7 T8 T9
T7 T8 T9
Figure 4.95 - Tool life (VC ≥ 0.3 mm or VN ≥ 0.6 mm) recorded when machining Ti-6Al-4V
alloy with different CBN/PCBN tools (T7, T8 and T9) grades with conventional coolant flow
(CCF), high coolant pressures of 11 MPa and 20.3 MPa at various cutting speed conditions.
206
4.4.2 Tool wear when machining Ti-6Al-4V alloy with different grades of CBN/PCBN
tools
Figure 4.96 shows tool wear rates when machining Ti-6Al-4V alloy with different
CBN/PCBN tool (T7,T8,T9) inserts using conventional coolant flow and high pressure
coolant supplies of 11 MPa and 20.3 MPa and at various cutting speeds. It can be seen from
the graph that the lower wear rate with regular wear pattern, increasing with increasing cutting
speed, was observed when machining with the lower (50 vol.%) CBN content T7 inserts,
unlike machining with PCBN (T8 and T9) tool grades, which gave higher wear rates. Nose
wear and notching were the dominant failure modes for T7 inserts (Figures 4.97-4.99)
whereas notching and chipping were the dominant failure modes for T8 and T9 tools (Figure
4.100). Note that machining trials with CBN/PCBN tool grades using conventional coolant
flow were carried out only at a cutting speed of 150 m min
-1
. Machining under conventional
coolant flow gave higher tool wear rates than with high pressure coolant supplies. The least
wear rate was recorded when machining with the lower (50 vol.%) CBN content T7 insert
using the highest pressure of 20.3 MPa at a cutting speed 150 m min
-1
whereas the highest
wear rate was recorded when machining with T8 tool grade (90 vol.%) PCBN content using
conventional coolant flow, Figure 4.96. Figure 4.96 also shows that increasing coolant
pressure from 11 to 20.3 MPa did not provide any reduction in wear rate when machining
with both T7 andT9 inserts at the conditions investigated. Rapid increase in notch wear rate
was recorded when machining with T8 insert using the highest coolant supply pressure of
20.3 MPa at high speeds in excess of 200 m min
-1
.
Figures 4.97–4.104 show worn CBN/PCBN tools after finish turning of Ti-6Al-4V alloy
under various machining conditions. All of these figures clearly show that notching occurred
in all the CBN/PCBN tools tested under all the conditions investigated. Chipping of the
cutting edge also occurred in most cutting conditions as illustrated in Figures 4.97-4.99,
4.101-4.104. Figures 4.97-4.99 show that, although tool wear rate decreased, machining of Ti-
6Al-4V alloy with T7 insert under high pressure coolant supplies did not prevent chipping
occurring in comparison with conventional coolant flow.
207
0
2
4
6
8
10
12
14
150 200 250
Cutting speed (m/min)
Wear rate (mm/min)
T7 (CCF) T7 (11MPa) T7 (20.3 MPa
)
T8 (CCF) T8 (11MPa) T8 (20.3 MPa
)
T9 (CCF) T9 (11MPa) T9 (20.3 MPa
)
Figure 4.96 - Wear rate curves of different CBN/PCBN tools when machining Ti-6Al-4V
alloy with conventional coolant flow and high pressures coolant supplies at various speed
conditions.
(a) Tool life < 1.6 min (b) Tool life < 0.5 min
Figure 4.97 - Worn PCBN 10 (T7) inserts after machining Ti-6Al-4V alloy using
conventional coolant supply at a speed of (a) 150 m min
-1
and (b) 200 m min
-1
.
Rake face
Notching
Chipping
208
(a) Tool life < 2.7 min (b) Tool life < 1 min
Figure 4.98 - Worn PCBN 10 (T7) inserts after machining Ti-6Al-4V alloy with 11 MPa
coolant pressure at a speed of (a) 150 m min
-1
and (b) 250 m min
-1
.
(a) Tool life < 4 min (b) Tool life < 0.6 min
Figure 4.99 - Worn PCBN 10 (T7) inserts after machining Ti-6Al-4V alloy with 20.3 MPa
coolant pressure at a speed of (a) 150 m min
-1
and (b) 250 m min
-1
.
Rake face Chipping
Rake face
Chipping
Chipping
209
Tool life < 0.2 min
Figure 4.100 - Worn PCBN 300 (T8) inserts after machining Ti-6Al-4V alloy using
conventional coolant supply at a speed of 150 m min
-1
.
(a) Tool life < 0.5 min (b) Close-up view
Figure 4.101 - (a) Worn PCBN 300 (T8) insert after machining Ti-6Al-4V alloy with 11 MPa
coolant pressure at a speed of 150 m min
-1
and (b) enlarged section of worn surface.
Notching
Rake faceChipping
Adhered work material
Notching
Work
material
210
(a) Tool life < 0.3 min (b) Tool life < 0.3 min
Figure 4.102 - Worn PCBN 300 (T8) inserts after machining Ti-6Al-4V alloy with 20.3 MPa
coolant pressure at a speed of (a) 200 m min
-1
and (b) 250 m min
-1
.
(a) Tool life < 0.2 min (b) Tool life < 0.3 min
Figure 4.103 - Worn PCBN 300-P (T9) inserts after machining Ti-6Al-4V alloy with
conventional coolant supply at a speed of (a) 150 m min
-1
and (b) with 11 MPa coolant
pressure at a speed of 250 m min
-1
.
(a) Tool life < 0.3 min (b) Tool life < 0.5 min
Figure 4.104 - Worn PCBN 300-P (T9) inserts after machining Ti-6Al-4V alloy with a
20.3 MPa coolant pressure at a speed of (a) 150 m min
-1
and (b) 200 m min
-1
.
Notching
Flank face
Flank face
Notching
Rake face
Notching
Rake face
Flank wear
Notching
Flank wear
Rake faceChipping
211
4.4.3 Component forces
Figures 4.105 and 4.106 show variations in cutting and feed forces, respectively,
recorded at the beginning of cut when machining Ti-6Al-4V alloy with CBN (T7) and PCBN
(T8,T9) tools. Figure 4.105 shows that cutting forces generally increase with increase in
cutting speed up to 200 m min
-1
when machining with all the coolant supply pressures
investigated. Higher cutting forces were generated when machining with all the CBN/PCBN
grades using 11 MPa coolant supply at a cutting speed of 200 m min
-1
. Figure 4.105 also
shows that relatively lower cutting forces were recorded when machining with high pressure
coolant supplies at a speed of 250 m min
-1
. This may be attributed to the severe notching and
excessive chipping of the cutting edges as a result of higher temperatures generated at higher
speed conditions.
50
100
150
200
250
150 200 250
Cutting speed (m/min)
Cutting force, Fc (N)
T7 (CCF) T7 (11MPa) T7 (20.3 MPa)
T8 (CCF) T8 (11MPa) T8 (20.3 MPa)
T9 (CCF) T9 (11MPa) T9 (20.3 MPa)
Figure 4.105 - Cutting forces (Fc) recorded at the beginning of cut when machining Ti-6Al-
4V alloy with different CBN/PCBN inserts at various cutting conditions.
Figure 4.106 shows that recorded feed forces were slightly lower than cutting forces at
all the cutting conditions investigated. Lower feed forces were generated when machining
with conventional coolant flow at the lower cutting speed of 150 m min
-1
like the cutting
forces. Higher feed forces were generated when machining with all the CBN/PCBN insert
grades with 11 MPa coolant supply at a cutting speed of 200 m min
-1
. Feed forces generally
increased with increase in cutting speed up to 200 m min
-1
when machining with all the
coolant supply pressures employed.
212
50
100
150
200
250
150 200 250
Cutting speed (m/min)
Feed force, Ff (N)
'
T7 (CCF) T7 (11MPa) T7 (20.3 MPa)
T8 (CCF) T8 (11MPa) T8 (20.3 MPa)
T9 (CCF) T9 (11MPa) T9 (20.3 MPa)
Figure 4.106 - Feed forces (Ff) recorded at the beginning of cut when machining Ti-6Al-4V
alloy with different CBN/PCBN inserts at various cutting conditions.
4.4.4 Surfaces roughness and runout values
Figure 4.107 illustrates the surface roughness values recorded at the end of cut when
machining Ti-6Al-4V alloy with CBN/PCBN inserts under various cutting conditions. Higher
surface roughness values were recorded when machining with the lower (50 vol.%) CBN
content (T7) insert at speeds of 200 m min
-1
and 250 m min
-1
using 11 MPa coolant pressure.
Surface roughness values increased with increase in cutting speed when machining with T7
tools using 11 MPa coolant pressure. Values recorded when machining with T8 and T9 tools
were relatively constant in all the cutting speeds investigated. Figure 4.107 also shows that
increase in coolant pressure has negligible effect on the surface roughness values when
machining with T8 and T9 tool grades. They varied between 1 and 2
µm.
In general, low runout values well below the stipulated rejection criterion of 100
µm,
were recorded when machining Ti-6Al-4V alloy with all the CBN/PCBN tools at the cutting
conditions investigated. Figure 4.108 are plots of runout values recorded when machining Ti-
6Al-4V alloy with all the CBN/PCBN tools under various coolant supply pressures at a speed
of 150 m min
-1
. The curves show that runout values vary between 5 and 18 µm. Additionally,
runout values generally decreased when machining at high pressure coolant supplies. Lower
213
runout values were recorded with the lower (50 vol.%) CBN content (T7) grade than with T8
and T9 insert grades.
0
1
2
3
4
5
6
150 200 250
Cutting speed (m/min)
Surface roughnes
s
Ra ( m)
T7 (CCF) T7 (11MPa) T7 (20.3 MPa)
T8 (CCF) T8 (11MPa) T8 (20.3 MPa)
T9 (CCF) T9 (11MPa) T9 (20.3 MPa)
Figure 4.107 - Surface roughness values recorded at the beginning of cut when machining
Ti-6Al-4V alloy with CBN/PCBN inserts at various cutting conditions.
0
5
10
15
20
T7 T8 T9
Cutting tool
Runout (
Ξ
m)
CCF HP 11MPa HP 20.3 MPa
Figure 4.108 – Runout variation recorded at end of cut when machining Ti-6Al-4V alloy with
CBN/PCBN inserts using conventional coolant flow and high coolant supply pressures at a
speed of 150 m min
-1
.
214
4.4.5 Surface hardness and subsurface alteration
Figures 4.109-4.111 are plots of variation of microhardness values with distance below
the machined surface of sectioned samples after machining with CBN/PCBN (T7, T8 and T9)
tools with various coolant supply pressures at a cutting speed of 150 m min
-1
. The plots
generally show a regular hardness pattern, i.e. evidence of softening of machined surface up
to about 0.15 mm below the top machined surfaces and a uniform distribution of hardness
values around the minimum and maximum values of the hardness prior to machining. The
plots also suggest that, in general, hardness depth of the machined surface decreased with
increasing coolant pressure. Lower microhardness values were recorded when machining with
T9 grade under all the coolant supplies.
200
250
300
350
400
450
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
CCF Min._Value (341 HV)
HP 11MPa Max._Value (363 HV)
HP 20.3MPa
Figure 4.109 - Hardness variation after machining Ti-6Al-4V alloy with PCBN 10 (T7) tools
with conventional coolant flow and high coolant supply pressures at a speed of 150 m min
-1
.
200
250
300
350
400
450
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness (HV
100
)
CCF Min._Value (341 HV)
HP 11MPa Max._Value (363 HV)
HP 20.3MPa
Figure 4.110 - Hardness variation after machining Ti-6Al-4V alloy with PCBN 300 (T8) tools
with conventional coolant flow and high coolant supply pressures at a speed of 150 m min
-1
.
215
200
250
300
350
400
450
0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60
Distance below machined surface (mm)
Microhardness
(HV
100
)
CCF Min._Value (341 HV)
HP 11MPa Max._Value (363 HV)
HP 20.3MPa
Figure 4.111 - Hardness variation after machining Ti-6Al-4V alloy with PCBN 300-P (T9)
tools with conventional coolant flow and high coolant supply pressures at a speed of
150 m min
-1
.
The microstructure of the etched machined surfaces after machining with CBN/PCBN
(T7, T8 and T9) tools using various coolant supply pressures at a cutting speed of 150 m min
-1
are shown at Figures 4.112-4.114, respectively. These figures exhibit similar characteristics.
The well defined grain boundaries show that there is no microstructure alteration in the
machined subsurface of the machined surface after machining under the cutting conditions
investigated.
(a) Tool life <1.6 min (b) Tool life < 4 min
Figure 4.112 - Microstructure of Ti-6Al-4V alloy after machining with PCBN 10 (T7) tools
with (a) conventional coolant flow and (b) high coolant pressure of 20.3 MPa at a cutting
speed of 150 m min
-1
.
216
(a) Tool life < 0.5 min (b) Tool life < 0.3 min
Figure 4.113 - Microstructure of Ti-6Al-4V alloy after machining with PCBN 300 (T8) tools
at (a) 11 MPa and (b) 20.3 MPa coolant pressure at a cutting speed of 150 m min
-1
.
(a) Tool life < 0.2 min (b) Tool life < 0.3 min
Figure 4.114 - Microstructure of Ti-6Al-4V alloy after machining with PCBN 300-P (T9)
tools at (a) 11 MPa and (b) 20.3 MPa coolant pressure at a cutting speed of 150 m min
-1
.
4.4.6 Chips shapes
Figures 4.115 (a)–(g) show different chips generated when machining with CBN/PCBN
tools under various coolant supply pressures at a cutting speed of 150 m min
-1
. These figures
show that most of the chips produced during machining Ti-6Al-4V alloy are helical, short and
continuous type. These figures also suggest that increasing coolant pressure provided a
negligible effect on the nature of swarf produced when machining Ti-6Al-4V alloy with the
CBN/PCBN tools. However, machining with the lower (50 vol.%) CBN content (T7) insert
217
under high pressure coolant supplies produced relatively smaller segmented chips (Figures
4.115 (b)-(c)) compared to T8 and T9 insert grades.
(a) T7 (conventional)
(b) T7 (11 MPa)
(c) T7 (20.3 MPa)
(d) T8 (conventional)
V = 150 m min
-1
(e) T8 (11 MPa)
V = 150 m min
-1
(f) T8 (20.3 MPa)
V = 150 m min
-1
(g) T9 (conventional)
Figure 4.115 - Chips generated when machining Ti-6Al-4V alloy with CBN/PCBN tools with
various coolant supplies at a cutting speed of 150 m min
-1
.
4.5 Machining of Ti-6Al-4V alloy with whisker reinforced ceramic cutting tools under
various machining environments
4.5.1. Wear rate and tool life
Figures 4.116 and 4.117 show nose wear rate and tool life, respectively, recorded when
machining Ti-6Al-4V alloy with silicon carbide (SiCw) whisker reinforced alumina ceramic -
T10 (rhomboid-shaped) and T11 (square-shaped) inserts at various cutting environments.
218
Note that machining with SiCw ceramic (T10 rhomboid-shaped) inserts with conventional
coolant flow was carried out only at a speed of 140 m min
-1
while machining with SiCw
ceramic (T11 square-shaped) inserts were carried out only with conventional coolant flow
(CCF) at speeds of 110, 130 and 200 m min
-1
. Increase in cutting speed generally accelerated
tool wear with both T10 and T11 grades of ceramic inserts in all the environments tested,
consequently reducing tool life due to a reduction in tool-chip and tool-workpiece contact
length and the consequent increase in both normal and shear stresses at the tool tip
(GORCZYCA, 1987).
0.0
1.0
2.0
3.0
4.0
5.0
110 130 140 200
Cutting speed (m/min)
Nose wear rate (mm/min) '
T10 (Argon) T10 (11MPa)
T10 (CCF) T10 (20.3 MPa)
T11 (CCF)
Figure 4.116 - Nose wear curves of silicon carbide (SiCw) whisker reinforced alumina
ceramic - rhomboid-shaped (T10) and square-shaped (T11) - inserts after machining Ti-6Al-
4V at various cutting conditions.
Figure 4.116 also shows that machining with T10 inserts in the presence of argon gave
the highest nose wear rate in all the cutting speeds investigated, hence the worst performance
in terms of tool life as illustrated in Figure 4.117 due probably to poor thermal conductivity of
argon which tends to accelerate tool wear. Additionally, the poor thermal conductivity of
argon can only prevent combustion from taking place during machining and most of the heat
generated tends to concentrate at the cutting interface which further accelerates tool wear
during machining. It is clear from Figure 4.117 that very low tool life was obtained when
machining Ti-6Al-4V alloy with SiCw ceramic (T10 and T11 grades) inserts in all the
conditions investigated relative to other cutting tools employed in this study. The longest tool
life of 2.7 min was obtained when machining with T10 insert using coolant supply pressure of
11 MPa and at a speed of 140 m min
-1
. Figures 4.116 and 4.117 also show that machining
219
under high pressure coolant supplies of 11 MPa and 20.3 MPa resulted to reduced tool wear
rate and consequently gave marginal improvement in tool life of T10 grade compared with
argon and conventional coolant supply. Machining with 20.3 MPa coolant pressure gave
slightly lower tool life than when machining with 11MPa coolant pressure. This can be
attributed to the physical phenomenon observed when the critical coolant pressure is exceeded
(PIGOTT; COLWELL, 1952). The optimum coolant pressure appears to have a relationship
with the total heat generated during machining (NAGPAL; SHARMA, 1973). It can also be
seen from Figures 4.116 and 4.117 that machining Ti-6Al-4V alloy with square-shaped
ceramic (T11) inserts exhibited inferior performance in terms of tool life than rhomboid-
shaped ceramic (T10) inserts in all the conditions investigated.
0.0
0.5
1.0
1.5
2.0
2.5
3.0
110 130 140 200 400 500
Cutting speed (m/min)
Tool life (min)
T10 (CCF) T10 (Argon)
T10 (11 MPa) T10 (20.3 MPa)
T11 (CCF)
Figure 4.117 - Tool life (nose wear, VC
≥ 0.3 mm or VN ≥ 0.6 mm) recorded when
machining Ti-6Al-4V alloy with silicon carbide (SiCw) whisker reinforced alumina ceramic -
rhomboid-shaped (T10) and square-shaped (T11) - inserts at various cutting conditions.
Ceramic tools generally exhibit lower fracture toughness than carbides and PCD tools in
addition to poor thermal and mechanical shock resistance. Additionally, ceramic tools have
high reactivity with titanium alloys. Typical wear patterns observed when machining titanium
alloys with ceramic cutting tools are notching and chipping of the cutting edge. Figures 4.118
(a)-(f) and 4.119 (a)-(b) are selected micrographs of worn SiCw reinforced alumina ceramic
tools, T10 and T11 grades respectively, after machining Ti-6Al-4V alloy under various
cutting conditions. Severe nose wear rate is the dominant failure mode observed when
machining with SiCw ceramic tools, especially with T10 insert at lower speed of 140 m min
-1
220
and with T11 insert at lower speed of 130 m min
-1
,
Figures 4.118 and 4.119, respectively.
Grooves can be clearly seen on the Figures 4.118 (a)-(f) and 4.119 (a), suggesting that SiCw
ceramics tools experienced mechanically related wear mechanism(s) mainly on the flank face,
illustrated as parallel ridges on flank faces. Severe chipping was also observed after
machining with T11 tool at cutting speed of 200 m min
-1
, as illustrated in Figure 4.119 (b).
221
(a) Tool life < 1.9 min (b) Tool life < 2.7 min
(c) Tool life < 0.4 min (d) Tool life < 1.9 min
(e) Tool life < 0.6 min (f) Tool life < 0.1 min
Figure 4.118 - Wear observed on rhomboid-shaped SiCw alumina ceramic (T10 grade) insert
after machining Ti-6Al-4V alloy with conventional coolant supply at a speed of (a) 140 m
min
-1
; coolant pressure of 11 MPa at speeds of (b) 140 m min
-1
and (c) 400 m min
-1
, coolant
pressure of 20.3 MPa at a speed of (d) 140 m min
-1
, and in an argon enriched environment at
speeds of (e) 200 m min
-1
and (f) 400 m min
-1
.
Severe
nose wear
Rake face
Severe
nose wear
Rake face
222
(a) Tool life < 0.8 min (b) Tool life < 0.4 min
Figure 4.119 - Wear observed on square-shaped SiCw alumina ceramic (T11 grade) insert
after machining Ti-6Al-4V alloy with conventional coolant supply at speeds of
(a): 130 m min
-1
and (b): 200 m min
-1
.
4.5.2 Component forces
Figures 4.120 and 4.121 are plots of cutting forces and feed forces, respectively,
recorded at the beginning of cut when machining Ti-6Al-4V alloy with SiCw whisker
reinforced alumina ceramic (T10 and T11) inserts at various cutting speeds and under various
machining environments. Note that cutting and feed forces generated with T10 inserts were
recorded only in presence of argon and with high pressure coolant supplies because negligible
tool life was generated when machining with conventional coolant flow. This machining time
was not enough to record component forces with the methodology employed. It can be seen
that recorded cutting forces are higher than feed forces, as expected. Higher cutting forces
were generated using T10 insert when machining with high pressure coolant supplies due
probably to higher wear rates. Figure 4.120 also shows that cutting forces recorded with T10
insert generally decreased with increasing cutting speed up to a speed of 400 m min
-1
in all the
machining environments, unlike when machining with T11 tool grade with conventional
coolant flow. Figure 4.121 also shows that higher feed forces were generated when machining
with high pressure coolant supplies. Feed forces generally increased with increasing cutting
speed when machining with high pressure coolant supplies and reduced with increase speed
when machining in argon environment. The lowest feed forces were recorded when
machining in presence of argon.
Chipping
Rake face
Severe nose
wear
Rake face
Grooves
223
0
60
120
180
240
300
360
420
110 130 140 200 400 500
Cutting speed (m/min)
Cutting force, Fc (N
)
T10 (Argon) T10 (11MPa)
T10 (20.3 MPa) T11 (CCF)
Figure 4.120 - Cutting forces (F
c) recorded at the beginning of cut when machining Ti-6Al-
4V alloy with SiCw alumina ceramic (T10 and T11) inserts at various cutting conditions.
0
60
120
180
240
300
360
420
110 130 140 200 400 500
Cutting speed (m/min)
Feed force, Ff (N)
'
T10 (Argon) T10 (11MPa)
T10 (20.3 MPa) T11 (CCF)
Figure 4.121 - Feed forces (Ff) recorded at the beginning of cut when machining Ti-6Al-4V
alloy with SiCw alumina ceramic inserts (T10 and T11) at various cutting conditions.
4.5.3 Surfaces roughness
Figure 4.122 shows recorded surface roughness values when machining Ti-6Al-4V
alloy with
SiCw whisker reinforced alumina ceramic (T10 and T11) inserts at various cutting
speeds and with T10 inserts with various machining environments. Note that surface
roughness obtained when machining under conventional coolant flow and in presence of
argon were recorded only at cutting speeds of 140 m min
-1
and 200 m min
-1
, respectively,
because at other cutting conditions the areas of machined surfaces generated were not enough
224
to measure surface roughness. It can be seen that surface roughness values recorded when
machining with T10 inserts using coolant supply pressures of 11 MPa and 20.3 MPa generally
increase with increase in cutting speed up to 400 m min
-1
. Surface roughness values also
increased with increasing in cutting speed when machining with T11 insert using
conventional coolant flow. Curves in Figure 4.122 also shows that higher surface roughness
values (between 2.1 µm and 5.5 µm) were recorded with conventional coolant flow and high
pressure coolant supplies of 11 MPa and 20.3 MPa which are well above the stipulated
rejection criterion of 1.6
µm.
0.0
1.0
2.0
3.0
4.0
5.0
6.0
110 130 140 200 400 500
Cutting speed (m/min)
Surface roughnes
s
Ra ( m)
T10 (Argon) T10 (20.3 MPa)
T10 (CCF) T11 (CCF)
T10 (11MPa)
Figure 4.122 - Surface roughness values recorded at the beginning of cut when machining Ti-
6Al-4V alloy with SiCw alumina ceramic inserts (T10 and T11) at various cutting conditions.
4.5.4 Surface hardness and subsurface alterations
Figure 4.123 illustrates the variations of microhardness values recorded from the
machined surfaces, up to about 1.0 mm below the top surface, after machining with silicon
carbide (SiCw) whisker reinforced alumina (T10) ceramic tool using various machining
environments and at a cutting speed of 140 m min
-1
. The plots suggest a softening of the
machined surface up to a distance of the about 0.7 mm below the machined surface when
machining with conventional coolant supply. The plots also show evidence of hardening when
machining with the coolant pressure of 20.3 MPa.
225
250
300
350
400
450
0.00 0.20 0.40 0.60 0.80 1.00
Distance below machined surface (mm)
Microhardness (HV100
CCF Min._Value (341 HV)
HP 11MPa Max._Value (363 HV)
HP 20.3MPa
Figure 4.123 - Hardness variation after machining Ti-6Al-4V alloy with rhomboid-shaped
SiCw alumina ceramic insert (T10 grade) at various environments and at a speed of
140 m min
-1
.
Figures 4.124 (a)-(c) are microstructures of etched machined surfaces of Ti-6Al-4V
alloy after machining with silicon carbide (SiCw) whisker reinforced alumina (T10) ceramic
tool using conventional coolant flow, high pressure coolant supplies of 11 MPa and 20.3
MPa, respectively, at a cutting speed of 140 m min
-1
. These figures show well defined grain
boundaries without microstructure alteration in the machined subsurface of machined
surfaces.
226
(a) Tool life <1.9 min (b) Tool life < 2.7 min
(c) Tool life < 1.9 min
Figure 4.124 - Microstructure of Ti-6Al-4V alloy after machining with SiCw alumina ceramic
tool (T10 grade) under: (a) conventional coolant flow, (b) coolant pressure of 11 MPa and
(c) coolant pressure of 20.3 MPa at a cutting speed of 140 m min
-1
.
4.5.5 Chips shapes
Figures 4.125 (a)–(f) show the types of chips generated when machining with SiCw
alumina ceramic (T10 and T11) inserts at various cooling environments. Note that Figures
4.125 (a)-(c) show chips produced when machining with T10 insert at a cutting speed of 140
m min
-1
and Figures 4.125 (d)-(e) show chips produced when machining with T11 insert at
cutting speeds of 130 and 200 m min
-1
, respectively. Three types of chips were produced
when machining with various machining environments investigated. Machining with T10 tool
with conventional coolant flow and in presence of argon produced snarled type chips (Figures
4.125 (a) and (b), respectively) while either short continuous tubular chips and segmented
type chips were produced when machining with high pressure coolant supply of 11 MPa.
227
Snarled type chips were also produced when machining with T11 insert in all the conditions
investigated, as illustrated in Figures 4.125 (d)-(e).
(a)
T10 - conventional
V=140 m min
-1
(b) T10 - argon
V=140 m min
-1
(c) T10 - 11 MPa
V=140 m min
-1
(d) T11 - conventional
V = 130 m min
-1
(e) T11 - conventional
V = 200 m min
-1
Figure 4.125 - Chips generated when machining Ti-6Al-4V alloy with SiCw alumina ceramic
inserts: T10 grade at a cutting speed of 140 m min
-1
under: (a) conventional coolant flow, (b)
argon enriched environment, (c) coolant pressure of 11 MPa; T11 grade with conventional
coolant flow at cutting speeds of: (d) 130 m min
-1
and (e) 200 m min
-1
.
4.6 Machining of Ti-6Al-4V alloy with Nano-ceramic cutting tools
4.6.1. Wear rate and tool life
Figures 4.126 and 4.127 show notch wear rate and tool life recorded when machining
Ti-6Al-4V alloy with nano-grain size (T12 and T13) ceramic tools with conventional coolant
flow and at various cutting speeds. Increase in cutting speed generally accelerated tool wear,
consequently reducing tool life, as expected, due to a reduction in tool-chip and tool-
workpiece contact lengths and the consequent increase in both normal and shear stresses at
the tool tip. Machining with the Al
2
O
3
base T12 tools gave higher wear rate when machining
at a cutting speed of 110 m min
P
-1
. The Si
3
N
4
base nano-grain T13 tool gave the lowest tool
228
wear rate. T13 tool, however, gave better tool life at a speed of 130 m min
-1
. Figure 4.127
shows that the nano-grain size tools gave poor performance, in terms of tool life, compared to
other cutting tools (cemented carbides, PCD, and CBN/PCBN tools) when machining Ti-6Al-
4V alloy at the conditions employed. It is clear from the machining results that improved
properties of the nanoceramic tools have negligible effect on tool performance. This suggests
that tool performance is more chemically and/or thermally related.
0.0
1.0
2.0
110 130 200
Cutting speed (m/min)
Notch wear rat
e
(mm/min)
T12 T13
Figure 4.126 - Notch wear rate when machining Ti-6Al-4V alloy with T12 and T13 nano-
grain size ceramics tools, Al
2
O
3
and Si
3
N
4
base respectively, with conventional coolant flow
and at various speed conditions
0.0
0.5
1.0
1.5
2.0
110 130 200
Cutting speed (m/min)
Tool life (min
)
T12
T13
Figure 4.127 - Recorded tool life (VN ≥ 0.6 mm) when machining Ti-6Al-4V alloy with
nano-grain size ceramics tools (T12 and T13) with conventional coolant flow and at various
cutting speeds.
229
Figure 4.128 (a)-(d) are typical wear on the worn nano-grain size ceramic tools after
machining Ti-6Al-4V alloy using conventional coolant flow at various cutting speeds.
Examination of the worn cutting edges revealed irregular/unevenly worn rake faces (Figure
4.128 (a), (b) and (c). Notching which often lead to catastrophic tool failure occurred when
machining with the Al
2
O
3
base nano-grain T12 tool at speeds in excess of 110 m min
-1
(Figure
4.128). Severe chipping was also observed after machining with T12 inserts at cutting speed
of 200 m min
-1
, Figure 4.128 (b). This type of wear occurs in a purely random manner and
cannot be predicted. Notch wear at the depth of cut can be observed in Figures 4.128 (a) and
(d). These appear to be mainly formed by a type of fracture process.
(a) Tool life < 0.6 min
(b) Tool life < 0.5 min
(c) Tool life < 0.8 min
(d) Tool life < 0.6 min
Figure 4.128 - Wear observed on nano-ceramic tools after machining with Ti-6Al-4V alloy at
different cutting speeds: T12 (a: 130 m min
-1
), (b: 200 m min
-1
); T13 (c: 110 m min
-1
) and
(d: 200 m min
-1
).
Notching
Rake faceRake face
Chipping
Flank face
Severe
nose wea
r
Rake face
Chipping
230
4.6.2 Component forces
Figure 4.129 shows variation of component forces (cutting forces and feed forces) with
cutting speed up to 200 m min
-1
recorded after 30 seconds of machining time when machining
Ti-6Al-4V alloy with nano-grain size (T12 and T13) ceramic tools with conventional coolant
flow. It is clear from Figure 4.129 that cutting forces generated were higher than feed forces
when machining with both grades of ceramic tools in all the conditions investigated. The
component forces generated increased with increasing cutting speed. This may be attributed to
the very high wear rate of the ceramic tools (Figure 4.128). This tends to increase frictional
forces during machining and the consequent loss/blunting of the sharp cutting edge.
0
50
100
150
200
250
300
110 130 200
Cutting speed (m/min)
Component force
s
Fc, Ff (N)
Fc (T12) Fc (T13)
Ff (T12) Ff (T13)
Figure 4.129 - Component forces (cutting forces: Fc and feed forces: Ff) recorded at the
beginning of cut when machining Ti-6Al-4V alloy with T12 and T13 tools with conventional
coolant flow.
4.6.3 Surfaces roughness and runout values
Figure 4.130 and 4.131 are the surface roughness and runout values, respectively,
recorded when machining Ti-6Al-4V alloy with nano-grain size (T12 and T13) ceramic tools
under conventional coolant flow and at various cutting speeds. Figure 4.130 clearly shows
that surface roughness values recorded are above the stipulated rejection criterion of 1.6 µm
due to severe wear on the tools. Increase in cutting speed when machining with Si
B
3
BNB
4
B base
nano-grain T13 tool grade had a negligible effect on the surface finish generated. Figure 4.130
also shows that high surface roughness values of 12.5 µm and 7.5 µm were recorded when
machining with the alumina base nano-ceramic T12 tool at cutting speeds of 130 and 200 m
min
P
-1
, respectively.
231
.
0
2
4
6
8
10
12
14
110 130 200
Cutting speed (m/min)
Surface roughnes
s
Ra ( m)
T12 T13
4.130 - Surface roughness values at the beginning of cut when machining Ti-6Al-4V alloy
with nano-ceramic (T12 and T13) tools with conventional coolant flow.
Figure 4.131 is a plot of the runout variation with speed when machining with nano-
grain size (T12 and T13) ceramic tools under conventional coolant flow. It can be seen that
runout values increased with increasing cutting speed for both tool grades investigated. The
silicon nitride base (T13) nano-grain size tool gave the lowest runout value of 7µm while the
alumina base nano-grain (T12) tool grade gave the highest runout values, rising to 20 µm at
the higher speed conditions of 200 m min
-1
. It should be noted that machining with nano-grain
ceramic inserts gave the highest runout values related to other tool grades investigated.
0
7
14
21
110 130 200
Cutting speed (m/min)
Run-out ( m)
T12 T13
Figure 4.131 – Runout values recorded at the end of cut when machining Ti-6Al-4V alloy cut
with nano-ceramic (T12 and T13) tools with conventional coolant flow
232
4.6.4 Chips shapes
Figures 4.132 (a)–(d) show chips generated when machining with nano-grain size (T12
and T13) ceramic tools with conventional coolant flow and at various cutting speeds. These
figures show that machining with nano (T12 and T13) ceramic tool generally produced
snarled type chips in all the speed conditions investigated.
(a) T12
V = 110 m min
-1
(b) T12
V = 130 m min
-1
(c) T12
V = 200 m min
-1
(d) T13
V = 200 m min
-1
Figure 4.132 - Chips generated when machining Ti-6Al-4V alloy with nano-ceramic tools:
T12 (a): 110 m min
-1
, b: 130 m min
-1
and c: 200 m min
-1
); T13 (d: 200 m min
-1
)
CHAPTER V
DISCUSSIONS
5.1 Introduction
As commented previously in Section 1.1 the main objective of this project is to achieve
significant reduction in cost of manufacturing jet engines by selecting the best combination of
cutting tool; cutting environment; cutting conditions for machining Ti-6Al-4V alloy. This
chapter will discuss the experimental results obtained when machining Ti-6Al-4V alloy with
various cutting tools (carbides, PCD, CBN and ceramics tool materials) under various
machining conditions. Effects of cutting tool material, machining environment such as argon
and with cutting fluid delivered at different pressures as well as machining variable such as
the cutting speed on the tool life, surface finish and surface damage of the machined
workpiece were also investigated.
5.2 Tool performance when machining Ti-6Al-4V alloy with different grades of carbide,
PCD, CBN and ceramic tools
5.2.1 Carbides tools
It can be seen from Figures 4.2, 4.59, 4.95, 4.117, 4.127 that increase in cutting speed
generally decreased tool life of all the cutting tool materials (T1-T13) investigated in this
thesis, as expected, due to a reduction in tool-chip and tool-workpiece contact lengths and the
consequent increase in both normal and shear stresses at the tool tip (GORCZYCA, 1987).
Reduction in tool-chip and tool-workpiece contact areas also tends to concentrate the high
temperature generated to a relatively smaller area as well as shifting the highest temperature
234
closer to the cutting edge. This phenomenon combines with higher stresses acting at the
cutting edge to promote the softening of the cutting tool, and consequently accelerating wear
processes and lowering tool life.
Tool life results obtained when machining with cemented carbides (T1,T2,T3,T4) are
shown in Figure 4.2. T1 (uncoated grade) and T4 (coated grade) inserts, which have the same
substrate composition with slightly coarser grain size and higher hardness than T2 (uncoated
grade) and T3 (coated grade) inserts (Table 3.3), gave the best performance in terms of tool
life in all conditions investigated relative to T2 and T3 insert grades with the same substrate
composition, hardness and substrate grain size. This may be attributed to the substrate grain
size and chemical composition. The substrate grain size of T2 and T3 inserts is 0.68
µm,
which is smaller than the substrate grain size of T1 and T4 inserts (Table 3.3). As previously
stated, hot compressive strength of cemented carbides depends on a combination of the Co
binder concentration and the WC particle size (LEE (1981), DEARNLEY; GREARSON
(1986), SZESZULSKI; THANGARAJ; WEINNMANN (1990)). The binding metal (Co)
easily dissolves in titanium thus carbides containing only small amounts of cobalt should be
used. On other hand, a lower Co content leads to reduced toughness/rigidity of the cutting
edge. The most favourable compromise between a limited Co content and adequate rigidity of
the cutting edge for machining titanium is found in the carbides of the ISO K10 and K20
grades (DEARNLEY; GREARSON, 1986). Influence of grain sizes on wear resistance of
carbides tools was investigated by Mari and Gonseth (1993) that attributed the higher
resistance to deformation of carbide tool at higher temperature to the coarser grain size. In
other studies (JAWAID; CHE-HARON; ABDULLAH (1999), CHE-HARON (2001)) lower
flank wear rate of carbide tools, hence longer tool life, is achieved with grain size of 1.0
µm
compared to finer grain size of 0.68
µm. The higher wear rate of finer grain size tools was
attributed to the increased solubility of WC in titanium alloys as the surface area of tool
particles exposed to solution wear increased. The relatively superior performance of T1 and
T4 inserts compared to T2 and T3 inserts may also be attributed to presence of tantalum
carbide (TaC) in their composition. TaC addition generally improves crater wear resistance
and increases the hot hardness of the carbides, thus preventing plastic deformation of the
cutting edge when machining at higher speed conditions (BOOTHROYD; KNIGHT, 1989).
Figure 4.2 also shows that T1 tool outperformed T4 tool when machining with
conventional coolant flow and in the presence of argon at all the speed conditions
investigated. However, T4 insert generally outperformed T1 insert when machining under
235
high coolant supply pressures up to 20.3 MPa. The relatively inferior performance of T4 tool
when machining under conventional coolant flow and in presence of argon can be attributed
to the PVD multi-layer (TiAlN /TiN) coating of T4 (Table 3.3) which has its lubrication
property suppressed by chemical reactivity with Ti element of the workpiece at the cutting
conditions investigated. Titanium is very reactive with most cutting tool materials at elevated
temperatures (> 550ºC) during machining (KONIG (1979), HARTUNG; KRAMER (1982),
MOTONISHI et al. (1987), BROOKES; JAMES; NABHANI (1998)). Coatings are used in
cutting tools to provide improved lubrication at the tool-chip and tool-workpiece interfaces,
reduce friction and consequently lower temperatures at the cutting edge. TiN coating is a good
diffusion barrier (EDWARDS, 1993) and generally provides a very low friction ratio at the
tool the cutting edge and an excellent resistance to crater wear (LÓPEZ DE LACALLE et al.,
2000). However, Ti element present in coatings can react chemically with Ti in the workpiece
at high cutting temperatures during machining, thereby leading to higher wear rate due to
diffusion wear. Difusion wear is a mechanism mainly dependent on temperature rise and
involves exchange of atoms between the tool and work material at the tool-workpiece and
tool-chip interfaces. It was reported that coatings of TiN, TiC, Al
2
O
3
and HfN in carbide tools
were worn more rapidly than uncoated straight carbides by either dissolution-diffusion or
attrition wear mechanisms when machining Ti-6Al-4V alloy (KONIG, 1979). It has also been
reported that multilayer (TiC/TiCN/TiN) coatings of a mixed cemented carbide tool were
rapidly removed from cemented carbide substrate by adhesion wear mechanism taking place
at the cutting edge (NABHANI, 2001b). The high adhesive forces are likely to result in the
plucking of hard particles from the tool. Thus erosion of the coating layer/(s) exposes the
carbide substrate to extreme temperature at the tool cutting edge, consequently increasing the
crater wear depth.
Results obtained when machining with T4 tool relative to T1 tool show that machining
Ti-6Al-4V with high pressure coolant supplies can minimise chemical reaction of TiAlN
coatings with workpiece, thus providing improved lubrication at the tool-chip and tool-
workpiece interfaces. This phenomenon will minimise dissolution wear, thereby increasing
tool life. T2 (uncoated) and T3 (coated) tools, in general, exhibited similar performance in
terms of tool life at the cutting conditions investigated, except when machining with the
highest coolant pressure of 20.3 MPa and at speed of 110 m min
-1
, where over 65% increase
in tool life was achieved using coated T3 tool. Increasing coolant pressure up to 11 MPa
resulted to increased tool life of all the tools at the highest cutting speed of 130 m min
-1
.
236
However when machining with 20.3 MPa coolant supply pressure, tool life lowered slightly
relative to 11MPa pressure (Figure 4.2 and Table 4.1). Table 4.1 shows that tool life increased
with increasing coolant pressure for all the grades of carbide inserts employed. The ranking
order for carbide tools in terms of average gain in tool life relative to conventional coolant
flow was T4, T3, T2 and T1 (Table 4.1). These results show that the coolant pressure has a
significant effect on tool wear pattern hence recorded tool life when machining Ti-6Al-4V
alloy with carbide tools under finishing conditions.
When high pressure coolant technique is employed, a hydraulic wedge is created at the
tool-workpiece interface which allows the coolant jet to penetrate the interface deeply even at
very high speed machining conditions. This action reduces the tool-chip and tool-workpiece
contact length/area and also changes the chip flow direction. A resulting force acting upon the
chip is created from the pressure in the fluid wedge. Additionally, the temperature gradient is
reduced by penetration of the high-energy jet into the tool-chip interface and consequently
eliminating the seizure effect (MAZURKIEWICZ; KUBALA; CHOW, 1989). This in turn
provides adequate lubrication at the tool-chip interface with a significant reduction in friction
(EZUGWU; BONNEY; YAMANE, 2003). These combined with high velocity coolant flow
causes the breakage of the chips into very small segments. Because the tool-chip contact time
is shorter, the tool is less susceptible to dissolution wear caused by chemical reaction with
newly generated chips, especially titanium-alloy chips (LINDEKE; SCHOENIG; KHAN,
1991). Moreover, the efficiency of high pressure coolant technique is dependent upon the jet
pressure and cutting speed (PIGOTT; COLWELL (1952), DAHLMAN (2000)). It has been
established that machining at any cutting speed, the tool-chip interface temperature initially
decreases with an increase in jet pressure up to a critical pressure (or optimum jet-pressure)
above which it rises to a relatively constant value for pressures in excess of the critical
pressure (SHARMA; RICE; SALMON, 1971) as a result of the critical boiling action of the
coolant at the tool edge since it was possible to sweep the tool surface faster by the higher jet
speed. This phenomenon lowers the rate of boiling and reducing heat transfer which
presumably may vary with tool-work material combination as well as cutting conditions.
Additionally, cutting tools operate within a safety temperature zone with minimal tool wear
when machining at the critical coolant pressure as thermal stresses are kept to a minimum,
thereby prolonging tool life (MACHADO; WALLBANK, 1994). From this explanation, it is
clear from Figure 4.2 and Table 4.1 that 20.3 MPa coolant pressure is above the critical
pressure for machining Ti-6Al-4V alloy with all carbide tools employed at a cutting speed of
237
130 m min
-1
. Critical pressure phenomenon was also reported in the experiments carried out
by Nagpal and Sharma (1973), Kishi et al. (1975), Kovacevic; Cherukuthota; Mazurkiewicz
(1995), Dahlman (2000) and Bonney (2004). In all of these studies the authors reported that,
in general, when coolant pressure reached a certain optimum value, a further increasing in
coolant pressure was not found to be beneficial in further improving machining performance.
Bonney (2004) also reported that any increase in coolant pressure in certain cases does not
improve tool life when machining Inconel 718 at cutting speeds lower than 50 m min
-1
and a
feed rate o 0.25 mm rev
-1
. Kovacevic; Cherukuthota; Mazurkiewicz (1995) attributed this to
the fact that a high pressure water jet after penetrating to a certain depth into the tool-chip
interface is not capable of penetrating any deeper, hence overcoming the high contact
pressures at the tool-chip interface. This probably explains why proportional and/or further
improvement in tool life of carbide tools was not recorded at higher coolant pressure of
20.3 MPa when machining at speeds in excess of 120 m min
-1
.
Figure 4.2 also shows that machining in an argon enriched environment gave inferior
performance in terms of tool life relative to conventional and high pressure coolant supplies
when machining Ti-6Al-4V alloy with carbides (T1 and T4). Table 4.1 also shows average
reduction in tool life when machining in argon enriched environment relative to conventional
coolant supply where 47 % and 44% reduction were obtained with using T1 and T4 tools,
respectively. This shows that machining Ti-6Al-4V alloy with carbide tools in an argon
enriched environment is still not recommendable because cooling and/or lubrication
characteristics of argon gas may be suppressed at the cutting conditions investigated.
Additionally, the poor thermal conductivity of argon can only prevent combustion taking
place during machining and most of the heat generated tends to concentrate at the cutting
interface. This tends to further accelerate tool wear during machining.
5.2.2 PCD tools
Longer tool life was recorded when machining Ti-6Al-4V alloy with PCD (T5 and T6)
tools with high coolant supply pressures than with conventional coolant flow (Figure 4.59).
This Figure also shows that tool life of the all the PCD tools decreased with increase cutting
speed in all machining environments investigated, consequently providing significant
reduction in tool life when machining at higher speed conditions. PCD (T5 and T6) tool
grades exhibited exceptional performance in terms of tool life when machining Ti-6Al-4V
alloy relative to the carbide and other tool materials employed in all the machining conditions
238
investigated. It has been reported (NABHANI, 2001b) that PCD tools also exhibited
outstanding performance relative to carbides and CBN tools when machining titanium TA48
alloy at a cutting speed of 75 m min
-1
, a feed rate of 0.25 mm rev
-1
and a depth of cut of 1.0
mm under dry machining conditions. This was probably attributed to the formation of a
protective layer of titanium carbide on the rake face of the tool due to inter-diffusion of
titanium and carbon atoms at the cutting interface between the work and tool material during
machining. The saturated titanium carbide layer adheres to the cutting edge of the PCD tool to
prevent further diffusion of the tool material into the chip / work material, thus increasing tool
life. Although cemented carbides are also susceptible to the formation of this stable titanium
carbide interfacial layer, PCD inserts have higher hot hardness than carbides (Table 3.3) and
can therefore maintain consistent cutting edge with prolonged machining at elevated
temperature generated in higher speed conditions. Higher thermal conductivity of PCD inserts
relative to carbide tools is another relevant factor to be considered. Cutting tool materials with
higher thermal conductivity minimise thermal gradients and thermal shocks during machining
process because heat generated in the cutting zone can easily be dissipated into the chip and
the cooling environment hence the tool maintains it yield stress (EZUGWU; WANG (1997),
TRENT; WRIGHT (2000)). This will minimise/prevent the weakening of the tool bond
strength, thereby minimising tool wear rate and increasing tool life.
Longer tool life was recorded when machining with T5 tool using a medium coolant
supply pressure of 7 MPa relative to conventional coolant flow in all the cutting speeds
investigated and also relative to 11 MPa coolant pressure at cutting speeds of 175 and
230 m min
-1
. This irregular PCD´s tool life behaviour suggests that optimum cutting
conditions for machining titanium alloys will depend on the correct combination of several
variables such as coolant pressure, grain size of tool and cutting speed. Additionally, coolant
pressure employed has a significant effect on tool wear pattern and hence recorded tool life
when machining Ti-6Al-4V alloy with PCD tools under finishing conditions, similarly the
carbides tools. The best performance of larger grain size (T5) tool when machining Ti-6Al-4V
alloy can be achieved using the highest coolant pressure supply of 20.3 MPa at a cutting speed
range of 175-230 m min
-1
whereas the best performance of smaller grain size (T6) tool is
achieved when machining with coolant pressures lower than 20.3 MPa and at cutting speeds
up to 230 m min
-1
. High temperatures generated close to the cutting edge of the tool are the
principal reasons for the rapid tool wear commonly observed when machining titanium alloys
(TRENT; WRIGHT, 2000). About 80% of the heat generated is generally absorbed in the tool
239
when machining Ti-6Al-4V alloy while the machining of steel it is about 50% (KONIG,
1979). This suggests that the superior performance of T5 tools under the conditions
investigated is a result of their higher thermal conductivity compared to T6 tools (Table 3.3)
combined with higher heat removal rate provided by the higher coolant flow rate delivered to
the cutting zone at a coolant pressure supply of 20.3 MPa. On the other hand, the lower
thermal conductivity of T6 tool (459 Wm
-1
K
-1
against 540 Wm
-1
K
-1
) caused more heat to
develop at the cutting zone, thus assisting in the softening of the work material. The improved
performance of T6 tool may be attributed to its smaller grain size. It has been reported
(COOK; BOSSOM (2000), BAI et al. (2004)) that abrasion resistance of PCD tools and
consequently tool life increases with increase in grain size. However, this correlation is not
linear and tool life can be lowered if a grain size is increased beyond a certain value (COOK;
BOSSOM, 2000). This is dependent on the cutting conditions employed. Finally, it is
important to state that the outstanding performance, in terms of tool life, exhibited by PCD
tools relative to carbides is suppressed by their higher cost (over 8900% higher than cemented
carbides) (SECO TOOLS, 2002b).
Regard Costing and Productivity analysis, some benefits of using the High Coolant
Pressure Technology in Machining Ti-6Al-4V alloy with carbide and PCD tools were
achieved: results of research programme generated a generic technology that can be employed
by all industries engaged in superalloy machining. Therefore, the results from machining
technology developed in this research programme produced 50% reduction in Cycle times
demonstrated on JSF Blisk Turning; 5 fold reduction in consumables. Additionally, about
£ 1.5 million capital saving for JSF Factory and £ 1750.00 saving per machined part (based on
projected cost rate and reduction in cycle time).
5.2.3 CBN tools
Lower tool life was recorded when machining Ti-6Al-4V alloy with CBN (T7, T8 and
T9) inserts (Figure 4.95) relative to the carbide and PCD tools due probably to the high
reactivity of titanium alloys with CBN tools (VIGNEAU (1997). Tool life decreased with
increasing cutting speed when machining with the lower (50 vol.%) CBN content (T7 grade),
unlike when machining with T8 and T9 inserts with higher (90 vol.%) CBN content. T7 insert
gave the best performance in terms of tool life among all the CBN tools at all conditions
investigated. T8 and T9 inserts in general exhibited similar performance in terms of tool life
(generally below 1 min) at the conditions investigated. Additionally, the coating of the T9 tool
240
did not influence tool performance. The results from Figure 4.95 suggest that the relatively
good performance of T7 grade compared to T8 and T9 grades are associated with the lower
CBN content of the T7 substrate as well as the high temperature characteristics of the ceramic
matrix (Table 3.3). Superior performance in terms of tool life of low CBN content tools
relative to those of high CBN content tools when machining difficult-to-machine materials,
especially titanium alloys, is attributed to the ability of the lower CBN content grades to retain
their cutting edge for longer period (KOMANDURI; REED JR (1983), RICHARDS;
ASPINWALL (1989), HUANG; LIANG (2003)). Table 3.3 also shows that the higher CBN
contents, T8 and T9, inserts exhibit much higher thermal conductivity than T7 insert (138Wm
-
1
K
-1
against 44Wm
-1
K
-1
). This can enhance the weakening of the CBN-ceramic bond at higher
speed conditions, thereby accelerating tool failure by mechanically related modes such as
chipping and notching as illustrated in Figures 4.97-4.104.
Figure 4.95 also shows that tool life of all the CBN inserts generally increased with
increasing coolant pressure, especially when machining with the T7 tool at speeds of 150 and
250 m min
-1
, which are more sensitive to coolant pressure than others. Over 68% and 150%
improvement in tool life was recorded when machining with T7 tool at a speed of 150 m min
-1
with coolant pressures of 11 MPa and 20.3 MPa, respectively, relative to conventional coolant
flow. However, tool life decreased with increasing the coolant pressure when machining at
higher cutting speeds in excess of 150 m min
-1
. This tool life pattern can be attributed to the
physical phenomenon observed when the critical coolant pressure is exceeded (PIGOTT;
COLWELL, 1952). The optimum coolant pressure appears to have a relationship with the
total heat generated during machining. A major benefit of coolant delivery at higher pressures
is improved access of the coolant to the cutting interface and the subsequent reduction of the
cutting temperature, hereby minimising and/or completely eliminating thermally related wear
mechanisms (BONNEY, 2004).
5.2.4 Micron-grain ceramic tools
Lower tool life (< 2 min) was recorded when machining Ti-6Al-4V with various grades
(T10,T11,T12 and T13) of ceramics tools under various cutting conditions relative to
carbides and PCD tools (Figures 4.2, 4.59, 4.117 and 4.127). Ceramic tools generally exhibit
lower fracture toughness than carbides as well as poor thermal and mechanical shock
resistance. Poor performance of ceramics tools, in terms of wear rate and consequently very
low tool life can be attributed to their low thermal conductivity compared to other
241
commercially available cutting tool materials (Table 3.3) as well as to their relative low
fracture toughness and high reactivity with titanium alloys (YANG (1970), (KONIG (1979),
LEE (1981), KOMANDURI; REED JR (1983), LI; LOW (1994)). The reduction of hot
hardness at elevated temperatures conditions during machining lead to the weakening of the
inter-particle bond strength and the consequent acceleration of tool wear (NORTH, 1986).
Rhomboid-shaped ceramic (T10) inserts in general outperformed squared-shaped ceramic
(T11) inserts when machining at relatively low cutting speeds, (< 200 m min
-1
), using
conventional coolant flow (Figure 4.117). Note that machining with SiCw ceramic (T10)
insert under conventional coolant flow was carried out only at a speed of 140 m min
-1
.
Increase in cutting speed generally accelerated tool wear when machining with all ceramic
tool grades tested in all the environments investigated due to a reduction in tool-chip and tool-
workpiece contact length and the consequent increase in both normal and shear stresses at the
tool tip. The reduction in tool-chip and tool-workpiece contact areas also tend to concentrate
the high temperature generated to a relatively small area as well as shifting the highest
temperature closer to the cutting edge (GORCZYCA, 1987). Machining in presence of argon
also gave the highest nose wear rate in all the cutting speeds investigated, hence the worst
performance in terms of tool life relative to other cooling environments. This performance
may be attributed to poor thermal conductivity of argon which tends to accelerate tool wear.
From Figure 4.116 also can be seen that increase in coolant pressure up to 11 MPa lowered
tool wear rate, consequently improved tool life, especially at lower speed conditions. Similar
to the phenomenon observed when machining with carbides and PCD tools, machining with
coolant supply pressure of 20.3 MPa and at cutting speeds up to 400 m min
-1
gave lower tool
life compared to coolant pressure of 11MPa. This can be attributed to the physical
phenomenon observed when the critical coolant pressure is exceeded (PIGOTT; COLWELL,
1952), as previously stated. This explains why proportional and/or further improvement in
tool life was not recorded when machining Ti-6Al-4V alloy with rhomboid-shaped SiC (T10)
ceramic inserts at the highest coolant pressure of 20.3 MPa at the cutting conditions
investigated.
5.2.5 Nanoceramic tools
Al
2
O
3
base (T12) and Si
3
N
4
base (T13) nano-grain ceramic inserts gave the least
performance in terms of tool life when machining Ti-6Al-4V alloy under conventional
coolant flow (Figures 4.126 and 4.127) relative to cemented carbides, PCD, CBN and
242
micron-grain ceramic inserts employed in this study. Increase in cutting speed generally
lowered tool life, as expected, due to a reduction in tool-chip and tool-workpiece contact
lengths and the consequent increase in both normal and shear stresses at the tool tip. It is also
clear from the machining results (Figure 4.127) that improved properties of the nanoceramic
(T12,T13) insert have negligible effect on tool life relative to commercially available ceramic
tools used for machining titanium alloys. This, therefore, suggests that the tool performance
of nano-grain ceramic tools is more chemically and/or thermally related. The Si
3
N
4
base
nano-grain
B T13 grade gave better performance in terms of tool life and tool wear rate relative
to T12 and micron-grain size SiC
w
alumina (T11) tools with the same tool geometry, as
illustrated in Figure 4.117 and Table 3.3. Although the micron-grain size SiCw alumina (T11)
and Al
2
O
3
base nano-grain (T13) inserts have theoretically improved edge toughness desired
for efficient machining, their relatively poor performance can be associated with their higher
hardness and associated brittleness (Table 3.3). The ranking order for the performance of the
nanoceramics tools, in terms of tool life, when machining Ti-6Al-4V alloy is T13 and T12.
5.3 Tool failure modes and wear mechanisms when machining Ti-6Al-4V alloy with
different grades of carbides, PCD, CBN and ceramic tools
5.3.1 Carbide tools
Adhesion, diffusion-dissolution, attrition and plastic deformation wear mechanisms are
typical tool failures of cemented carbide tools (KONIG (1979), HARTUNG; KRAMER
(1982), MACHADO; WALLBANK (1990), EZUGWU et al. (2005)). Diffusion wear
mechanism generally occurs where high temperature exists at the cutting interfaces. Adhesion
wear mechanism frequently occurs when there is chemical affinity between the tool and the
work material. High temperature generated at the chip-tool interface is a relevant phenomenon
in relation to the tribological system behaviour existing during machining operation. Although
attrition wear mechanism is commonly associated to lower cutting speeds, it can also occur in
situations where intermittent contact exists at the chip-tool interface. Cemented carbide tools
can also fail by abrasion wear mechanism as a result of the chipped off hard particles
sandwiched between the tool flank face and the newly machined surface. As the tool rubs
over the machined surface, the particles plough through the tool removing material at the
flank face (BROOKES; JAMES; NABHANI (1998), TRENT; WRIGHT (2000). This type of
243
wear mechanism has been reported when turning a titanium base, Ti-6242, alloy under dry
condition with carbide tools (JAWAID; CHE-HARON; ABDULLAH, 1999). The poor
thermal conductivity of the titanium alloy workpiece encourages the development of
maximum temperature closer to the cutting edge than is the case with other metals and alloys.
Thus, the formation of a crater in this region undermines the integrity of that cutting edge
resulting to fracture and accelerated wear rate (BROOKES; JAMES; NABHANI, 1998).
Typical failure modes observed when machining Ti-6Al-4V alloy with all cemented
carbides tools under various cutting conditions are nose wear and flank wear. The flank wear
rate is lower than the nose wear rate when machining with all carbide (T1,T2,T3,T4) inserts.
Since nose wear is the predominant failure mode observed in all conditions investigated
(Figures 4.3-4.5), flank wear curves were omitted in this thesis. This is due to the fact that
initial flank wear developed in all conditions investigated during the machining process tend
to be displaced towards the nose region with prolonged machining (Figures 4.6, 4.7, 4.10-4.17
and 4.20) and also with increasing cutting speed due to the reduction in the chip-tool and tool-
workpiece contact lengths areas (SADIK; LINDSTRÖM, 1995).Concentration of shearing
forces at the nose region usually increases the frictional forces and temperature that accelerate
wear at the tool nose (BONNEY, 2004). Rapid increase in nose wear occurred when
machining with all carbide inserts at speeds in excess of 110 m min
-1
in the presence of argon
and under conventional coolant flow (Figure 4.3). This, therefore, suggests that high pressure
coolant supplies are responsible for reduction in temperature at the cutting interface and
consequently low wear rates during machining, thereby increasing tool life. In other words,
the ability to remove heat from cutting tool increases with increase in coolant pressure.
Machado (1990) reported that high pressure coolant system reduces the size of the heat source
(reduced contact length) but not the amount of heat generated, hence higher temperatures in a
smaller heat affected zone at the tip of the tool would be expected. However, the author
reported that the high pressure coolant jet is expected to cause dissipation of the heat more
efficiently from the heat affected zone. This suggests that a smaller flow zone produces the
same amount of heat. This heat is, however, conducted more efficiently, thereby lowering the
cutting temperature. The temperature gradient reduces by penetration of the high-energy jet
into the tool-chip interface and consequently eliminating the seizure effect
(MAZURKIEWICZ; KUBALA; CHOW, 1989), thereby providing adequate lubrication at the
tool-chip interface with a significant reduction in friction (EZUGWU; BONNEY; YAMANE,
2003). These combined with high velocity coolant flow causes breakage of the chips into very
244
small segments. Because the tool-chip contact time is shorter, the tool is less susceptible to
dissolution wear caused by chemical reaction with newly generated chips, especially titanium-
alloy chips (LINDEKE; SCHOENIG; KHAN, 1991).
Machining with carbide tools under finishing conditions showed steady increase in nose
wear rate with increasing cutting speed (Figure 4.3). Additionally, nose wear rate generally
decreases with increasing coolant pressure. This is attributed to the fact that high coolant
pressure technique reduces the tool-chip contact length/area and also changes the chip flow
direction (as stated previously), thereby providing adequate lubrication at the tool-chip
interface with a significant reduction in friction (EZUGWU; BONNEY; YAMANE, 2003). In
general, flank wear rate observed when machining under high coolant pressures was very low
compared with nose wear rate. The mechanism of wear that will predominate depends on
what is happening at the interface during machining. Titanium reacts with most cutting tool
materials at elevated temperatures (> 550ºC) during machining (KONIG (1979),
(HARTUNG; KRAMER (1982), MOTONISHI et al. (1987), BROOKES; JAMES;
NABHANI (1998)). At such high temperature conditions titanium atoms diffuse into the
carbide tool material and react chemically with carbon present in the tool to form an interlayer
of titanium carbide (TiC) (KONIG (1979), (HARTUNG; KRAMER (1982)) which bonds
strongly to both the tool and the chip to form a saturated seizure zone. This action, minimise
diffusion wear mechanism by halting the reaction. The separation of the welded junction
results in tool material being carried away by the fast flowing chip. Titanium has a strong
tendency to adhere to the tool material. Micrographs of worn inserts (Figures 4.6, 4.11 and
4.17) show that adhesion of material was more pronounced when machining with T1, T2 and
T4 inserts, respectively, using conventional coolant flow. This suggests that adhesion of work
material to the tool nose is minimised when machining with high coolant pressure supplies.
Grain sizes of a tool substrate also affect its wear resistance. Nose wear rates of T2 and
T3 inserts were higher than those obtained when machining T1 and T4 inserts in all the
conditions investigated (Figure 4.3). Micrographs of T2 and T3 inserts in Figures 4.11 and
4.14 (b) illustrate the irregular wear pattern observed along the flank and rake faces of these
tools compared to T1 (Figure 4.6 (b)) and T4 (Figure 4.17) inserts. Inferior performance of T2
and T3 tools may be attributed to their smaller substrate grain size (0.68
µm) compared to
substrate grain size of T1 and T4 tools (1
µm) (Table 3.3). Similar results were reported in
several studies attributed the higher wear rate of finer grain size tools to increased solubility
of WC in titanium alloys as the surface area of tool particles exposed to increased solution
245
wear have been reported (MARI; GONSETH (1993), JAWAID; CHE-HARON;
ABDULLAH (1999), CHE-HARON (2001)). Additionally, the presence of TaC in
composition of T1 and T4 tools improved their crater wear resistance and increased their hot
hardness, thus preventing plastic deformation of the cutting edge to take place during
machining at high speed conditions (BOOTHROYD; KNIGHT, 1989).
Evidence of abrasion wear mechanism is shown in Figure 4.12 (a) from the visible
groove marks on the worn tool surfaces of T2 tool grade with a coolant pressure of 11MPa
and speed of 110 m min
-1
and T3 tool with high coolant pressures of 11 and 20.3 MPa as
illustrated in Figures 4.15 and 4.16, respectively. Grooves are more pronounced in coated
carbide T3 insert what suggesting that the wear mechanism is mechanically related and may
be generated by hard particles of the tool that have been detached by attrition and rubbed
against the flank surface. The erosion of the coating layer exposes the carbide substrate to
extreme temperature at the tool cutting edge causing severe grooving and the cratering
(NABHANI, 2001a). A similar process of attrition and grooving wear can also develop on the
flank face, leading to deterioration in the machined surface. Ultimately, the combination of
the crater and flank wear undermines the entire cutting edge and causing pronounced chipping
and/or fracture (NABHANI, 2001b). It is important to note that this wear mechanism was not
occur when machining with coated carbide T4 inserts. Figures 4.7-4.9 and 4.19 show a typical
worn tool showing flank and rake face wears. The uniform flank wear observed may be due to
the low wear rate caused by temperature reduction at the cutting interface when machining
with high coolant pressures.
The high adhesive forces are likely to result in the plucking of hard particles from the
tool. Therefore, erosion of the coating layer (s) exposes the carbide substrate to extreme
temperature at the tool cutting edge causing its surface to become grooved and the increasing
crater wear. T3 (coated carbide CP200 grade) insert experienced more severe flank and nose
wears in all machining environments tested relative to other carbide inserts. Visible grooves
were observed on the worn T3 inserts (Figures 4.15 and 4.16 (a) and (b)). Figure 4.20 (b)
shows that coated carbide T4 tool experienced severe flank and crater wears very close to the
cutting edge, when machining in presence of argon at a speed of 120 m min
-1
. The high
chemical reactivity of the titanium alloys results in excessive tool wear during machining.
Figures 4.12 (a), 4.15 and 4.16 (b) also suggest that T2 and T3 tools also experienced
appreciable crater wear. This failure mode is generally associated with high temperature
generated at the chip-tool interface and occurs on the rake face of a tool due to a combination
246
of diffusion and adhesion as the chip moves over the rake face of the tool (AB SANDVIK
COROMANT (1994), TRENT; WRIGHT (2000)). It was reported that when machining Ti-
6Al-4V alloy with cemented carbides a shallow crater was initially formed on the rake face of
the tool and a low flank wear was produced along the whole extension of the depth of cut
(MACHADO, 1990).
Figures 4.10 and 4.20 (a) show worn tools after machining in the presence of argon with
uncoated (T1) and coated carbide (T4) inserts at a speed of 130 m min
-1
and 100 m min
-1
,
respectively. Examination of these inserts revealed smoothly worn tool with slight cracks and
adhesion of work material to the tool nose as a result of the intermittent contact between the
tool and the workpiece. This suggests that high temperatures were generated at the cutting
interface and that diffusion wear mechanism is dominant under such conditions. The highest
nose wear rate recorded when machining with T1 and T4 inserts in the presence of argon can
be attributed to the poor thermal conductivity of argon which can only prevent combustion
taking place during machining. Most of the heat generated tends to concentrate at the cutting
interface, thus weakening the strength of tool and accelerating the tool wear during
machining.
5.3.2 PCD tools
Nose wear was the predominant failure mode observed when machining Ti-Al-4V alloy
with PCD tools in all the conditions investigated. Nose wear rate increased with increase in
cutting speed and with prolong machining using both PCD tools under all the coolant supplies
employed. There is a steady increase in nose wear rates on both PCD (T5 and T6) tools at
cutting speeds up to 160 m min
-1
when machining with conventional coolant flow (Figure
4.60). This low wear pattern is observed at cutting speeds up to 250 m min
-1
when machining
with high pressure coolant supplies of 7, 11 and 20.3 MPa. It is clear that the coolant pressure
has a significant effect on tool wear and hence recorded tool life when machining Ti-6Al-4V
alloy with PCD tools under finishing conditions. Rapid increases in nose wear rate occurred
when machining with both PCD (T5 and T6) tools at speeds in excess of 160 m min
-1
with
conventional coolant flow. As observed when machining with carbide tools at a different
cutting speed range, this wear pattern shows the ability to remove heat from cutting tool
increases with increase in coolant pressure. The strength of the cutting tools is, therefore,
expected to increase if a significant reduction in the tool temperature can be achieved,
particularly at high temperature regions typical of those encountered when machining
247
superalloys (MACHADO, 1990). Reduction in temperature can reduce tool wear rate during
machining, hence providing longest tool life. Figure 4.60 shows that nose wear rate can be
reduced at least by 88% by employing high pressure coolant technique in machining of Ti-
6Al-4V alloy with PCD, T5, tool grade at the cutting conditions investigated. Similar
percentage reduction in nose wear rate was also observed when machining with T6 insert.
Figure 4.61 illustrates the curves of progress of the nose wear
versus machining time when
cutting with both PCD grades at a cutting speed of 175 m min
-1
. The concentration of the
shearing forces at the nose region causes increased frictional forces and temperature that leads
to high wear rate at the nose region of the tool (BONNEY, 2004). Machining with
conventional coolant flow always provided accelerated nose wear, hence the lower tool life,
than with high pressure coolant supplies at the conditions investigated.
Diffusion-dissolution, attrition and plastic deformation are typical wear mechanisms
observed when machining titanium alloys with PCD tools. High temperature generated at the
chip-tool interface combined with the chemical affinity between the PCD tool and titanium
alloys when machining at higher speed conditions lead to significant reduction in tool life.
Although attrition wear mechanism is commonly associated to lower cutting speeds, it can
also occur in situations where intermittent contact exists at the chip-tool interface. Segmented
chips are generated at both lower and higher speed conditions when machining titanium
alloys. This phenomenon can enhance the stick-slip process, removing from the rake and
flank faces minute amount of tool particles by plucking action, thus accelerating tool wear
(MACHADO (1990),(NABHANI (2001b), BONNEY (2004)). The plucking process tend to
be accelerated by the application of high pressure coolant supply which is capable of
removing any loose tool material from the rake face as well as eroding the brittle PCD tool
particles. It has been reported that diffusion is an important wear mechanism in responsible
for cratering and it also predominates at the flank face leading to the development of the
maximum flank wear on tool (TRENT, 1988b). Additionally, rates of diffusion at the
interface normally increase exponentially with temperature. Notch wear is caused by abrasion,
mainly at the depth of cut line, of hardened saw tooth like material during machining.
Notching can also be accelerated by high temperature generated at the tool edge during
machining and subsequent weakening of the bond strength of the tool material. According to
Bonney (2004), the initial notch formed during machining usually act as a highly stressed
point and tends to accelerate tool wear leading to fracture in the worse case.
248
Figures 4.62-4.68 show micrographs of the worn PCD tools under various machining
conditions. Tool surfaces that may have been subjected to stick-slip action during machining
show a characteristic rough texture as illustrated in Figures 4.63-4.65, 4.67 and 4.68. Figure
4.62 suggests that adhesion of work material to the tool nose occurred when machining with
PCD tools under conventional coolant flow. Adhesion of work material to the nose of tools
may be caused by the intermittent contact between the tool and the workpiece. Notching on
was observed when machining with T6 insert under conventional coolant flow at a cutting
speed 175 m min
-1
(Figure 4.67). This may be attributed to the interdiffusion of titanium and
carbon at elevated temperature conditions developed during machining (KONIG (1979,
HARTUNG; KRAMER (1982), NABHANI (2001b)). The diffusion process results in the
formation of titanium carbide layer on the tool, which prevents further diffusion. Although
worn rake face of tools employed when machining with high pressure coolant supplies show
evidence of a tiny layer of work material on the both rake and flank faces due to high
chemical affinity between titanium from the workpiece and carbon from the cutting tool
(Figures 4.63-4.65, 4.67 and 4.68), crater wear seems to be more pronounced under these
conditions than with conventional coolant flow. This suggests that crater wear increased when
machining under high pressure coolant supplies. It has been reported that crater wear of
carbide tools increased with the coolant jet power when machining Ti-6Al-4V alloy with
carbide tools (VIGNEAU, 1997). It has also been reported that under certain conditions,
titanium adheres to the tool and no relative sliding occurs at tool-chip interface. A boundary
layer of titanium forms at the interface, and the relative motion between the tool and chip is
generated internally by shear within the titanium chip material. This layer will quickly
become saturated with tool constituents, limiting the mass transport of tool constituents from
the tool surface (HARTUNG; KRAMER, 1982). This would explain why no wear occurred in
certain regions of the crater zone of some tools (Figures 4.62 (b), 4.63 (b) and 4.66). Evidence
of notching at depth of cut can be seen in the worn T6 tool grade after machining at a speed of
175 m min
-1
under conventional coolant flow (Figure 4.66). However, this type of wear was
absent when machining with high pressure coolant supplies.
5.3.3 CBN tools
Figure 4.96 shows that machining with CBN (T7,T8,T9) inserts gave higher wear rates,
hence worse overall performance than carbides and PCD tools. The lower (50 vol.%) CBN
content T7 tool grade gave the lowest nose wear rate among the CBN tools investigated.
249
Cutting speed plays an important role in the performance of CBN tools. Increase in cutting
speed leads to further increase in cutting temperature as well as increases in the intensity of
chemical interaction between the tool and the work materials which adversely affects tool life.
As the cutting temperature increases, seizure of the chip occurs everywhere on the tool face.
This forms an adherent layer which becomes saturated with tool particles and serves as a
diffusion boundary layer, thus reducing the rate of transport of tool materials into the chip and
consequently the wear rate. Since the diffusivity rate increase exponentially with temperature,
further increase in cutting speed beyond the speed for minimum wear produces rapid increase
in the wear rate.
Nose wear, notching, chipping and premature tool failure are the dominant failure
modes encountered when machining with CBN tools due to their brittle nature (Figures 4.97-
4.105). Premature failure of CBN tools are generally attributed to the fracturing of the
unsupported cutting edge caused by attrition wear (BHAUMIK; DIVAKAR; SINGH, 1995).
However, average flank wear was the dominant failure mode when dry machining titanium-
base, Ti-5Al-4Mo-(2-2.5)Sn-(6-7)Si, alloy with CBN tools at a cutting speed of 75 m min
-1
, a
feed rate of 0.25 mm rev
-1
and a depth of cut of 1.0 mm (NABHANI, 2001a). Kramer (1987)
reported that CBN tools retain their strength at temperatures in excess of 1100ºC while
cemented carbide tools encounter plastic deformation at this temperature range. This
temperature is also that in which titanium can react with nitrogen in the CBN insert when
interatomic diffusion of titanium and CBN accelerates significantly.
Cutting speed also plays an important role in tools performance of CBN tools. Increase
in cutting speed leads to further increase in cutting temperature as well as increases in the
intensity of chemical interaction between the tool and the work materials which adversely
affects tool life. As the cutting temperature increases, seizure of the chip occurs everywhere
on the tool face. This forms an adherent layer which becomes saturated with tool particles and
serves as a diffusion boundary layer, thus reducing the rate of transport of tool materials into
the chip and consequently the wear rate. Since the diffusivity rate increase exponentially with
temperature, further increase in cutting speed beyond the speed for minimum wear produces
rapid increase in the wear rate.
CBN insert grades encountered severe notching and chipping of the cutting edges with
increase in cutting speed up to 250 m min
-1
and the consequent increase in cutting
temperature. These tend to lower tool life, even when using high pressure coolant supplies
(Figures 4.97 (b), 4.98 (b), 4.99 (b), 4.102 (b), 4.103 (b) and 4.104 (b)). The slightly better
250
performance of T7 grade relative to T8 and T9 grades may be associated with the lower CBN
content of the T7 substrate as well as the high temperature characteristics of the ceramic
matrix. Table 3.3 shows that the higher CBN content (T8 and T9) tools exhibit much higher
thermal conductivity. This can enhance the weakening of the CBN-ceramic bond at higher
speed conditions, thereby accelerating tool failure by mechanically related modes such as
chipping and notching (Figures 4.97-4.104). Machining under conventional coolant flow gave
higher tool wear rates than with high pressure coolant supplies. Tool wear rate generally
decreased with increasing in coolant pressure up to 11 MPa, especially when machining with
T7 insert (Figure 4.96). However, it can be seen from this figure that increasing coolant
pressure from 11 to 20.3 MPa did not provide any appreciable reduction in wear rate when
machining with both T7 ant T9 tool grades at the conditions investigated. This can be
attributed to the physical phenomenon observed when the critical coolant pressure is exceeded
(PIGOTT; COLWELL, 1952). Additionally, high pressure coolant supplies in general did not
prevent chipping from taking place. In some cases, CBN tools experienced fracture of cutting
edge, especially when machining with T8 grade (Figure 4.102 (b)). Notching was the
predominant failure modes present in all the CBN tools tested under all the conditions
investigated, followed by chipping of the cutting edge as illustrated in Figures 4.97-4.100,
4.102-4.103 and 4.104 (b). Dearnley and Grearson (1986) reported that CBN tools gave
higher wear rate than carbides when machining Ti-6Al-4V alloy at speeds in excess of
100 m min
-1
and therefore they were not considered suitable for practical application. Other
study reported that wear in CBN tools was associated with localized chipping of the cutting
edge as a result of tool-tip oscillations (ZOYA; KRISHNAMURTHY, 2000).
Rapid increase in notch wear rate was recorded when machining with T8 tool grade
under the highest coolant supply pressure of 20.3 MPa at high speeds in excess of
200 m min
-1
. Examination of worn CBN (T8 grade) tools revealed adhesion of work material
to both the nose and flank faces of tool as a result of the intermittent contact between the tool
and the workpiece (Figure 4.101). This figure shows that the thickness of the layer of material
stuck on the rake face of higher (90 vol.%) CBN content T8 tool grade may reach huge
proportions, even when machining with high pressure coolant supply. If thick layers of work
material stuck where the worn area develops, small fragments of it will flow down the flank
surface and cause further damage to the tool by attrition (MACHADO, 1990).
251
5.3.4 Micron-grain ceramic tools
Wear resistance, chemical inertness and toughness are considered main factors affecting
the performance of cutting tools when machining aerospace alloys, especially nickel and
titanium alloys (SZESZULSKI; THANGARAJ; WEINNMANN, 1990). Any ceramic tool
with a potential application for titanium machining must require a good thermal shock
resistance (LEE, 1981). Ceramic tool materials generally have three properties which
distinguish them from cemented carbides: chemically inert (with ferrous materials), highly
resistant to abrasive wear and capable of superior heat dispersal during chip forming process
(WHITFIELD, 1988). Silicon carbide (SiCw) whisker reinforced alumina ceramic has silicon-
carbide whiskers added to a matrix of aluminium oxide that reinforces the hard and somewhat
brittle aluminium oxide, allowing it to better withstand mechanical stresses. The fracture
toughness of silicon carbide (SiCw) whisker reinforced alumina ceramic (for instance,
WG300 grade) is enhanced by the phenomenon of whisker “pull-out” where the whiskers
generally pull out during the fracture process. Additionally, under actual cutting conditions
where temperatures at the tool-chip interface may reach over 1000ºC, whisker reinforced
alumina ceramic will retain high strength and hardness well beyond the point at which a
cemented carbide material has softened, deformed or failed completely (SMITH, 1986).
Machining with both (T10 and T11) ceramic inserts gave high nose wear rates, hence lower
tool life relative to carbide, PCD and CBN inserts employed in this study (Figure 4.116).
Ceramic tools generally exhibit lower fracture toughness than carbides and PCD tools in
addition to poor thermal and mechanical shock resistance. Additionally, ceramic tools have
high reactivity with titanium alloys (KONIG (1979), LEE (1981), DEARNLEY;
GREARSON (1986)). This factor is responsible for the worst performance of ceramic tools
relative to other tools. Machining in presence of argon with T10 ceramic tool gave the highest
nose wear rate in all the cutting speeds investigated due probably to the poor thermal
conductivity of argon which tends to accelerate tool wear. Machining with high pressure
coolant supplies of 11 MPa and 20.3 MPa reduced tool wear rate and consequently gave
marginal improvement in tool life for T10 grade compared with using argon and conventional
coolant flow (Figure 4.116). This may be attributed to the fact that high pressure coolant
technique reduces the tool-chip contact length/area and also changes the chip flow direction
(as stated previously), thereby providing adequate lubrication at the tool-chip interface with a
significant reduction in friction (EZUGWU; BONNEY; YAMANE, 2003). High pressure
coolant technique may have reduced tool temperature due to its ability to remove heat from
252
tool, thereby strengthen them for longer period during machining. Reduction of the
temperature can reduce tool wear rate during machining, hence prolonging tool life.
Figures 4.118 and 4.119 are micrographs of worn SiCw reinforced alumina ceramic
tools, T10 and T11 grades respectively, after machining Ti-6Al-4V alloy under various
cutting conditions. Typical failure modes observed with ceramics tools when machining Ti-
6Al-4V alloy are severe notching, chipping and, in some cases, severe nose wear. The
rhomboid-shaped (T10) ceramic grade experienced severe abrasive type wear mechanism
(Figures 4.118) while the square-shaped ceramic (T11) inserts experienced severe abrasive
wear on both the rake and flank faces (Figures 4.119). Abrasion wear mechanism involves the
loss of tool material by hard particles sandwiched between the tool and workpiece material.
The particles could be dislodged tool materials, fragments of built-up-edge or hard carbides
and oxides already existing in the workpiece. The high compressive stresses acting at the tool-
workpiece interface keep the trapped hard particles between the machined surface and the
tool. As the tool rubs over the machined surface, the particles plough through the tool
removing material at the flank face. In some cases this phenomenon lead to chipping of
cutting edge and eventually to catastrophic tool failure observed when machining at speeds in
excess of 140 m min
-1
as shown in Figures 4.118 (c) and (f) and 4.119 (b). This type of wear
occurs purely on a random manner and cannot be predicted. It has been reported that abrasive
wear mechanism in ceramic tools can also be caused by plastic flow of the tool material due
to the high temperature conditions encountered at the cutting interface (NORTH (1986),
DEARNLEY; GREARSON (1986), LI; LOW (1994), GATTO; IULIANO (1997)).
According to Bonney (2004), the laminar chip flow abrades the tool by stretching the
contacting plasticised tool layer until necking. This type of abrasion wear generally is
recognized by the presence of grooved surfaces as parallel ridges can be seen on the flank
face of tool (Figures 4.118 (a), (b), (d), (e) and 4.119 (a)). The formation of ridges on the
worn flank face is attributed to seizure between the workpiece and the tool flank face. This
type of ridges was also observed on ceramic worn tools when machining Inconel 718
(BONNEY, 2004). Although ceramic tools have high hardness at ambient temperature, plastic
behaviour of ceramics is possible at high temperature conditions when hydrostatic conditions
occur. This unusual thermo-mechanical property is exhibited by nano-ceramics and is referred
to superplasticity, defined as the ability of polycrystalline solids to exhibit exceptionally large
elongation in tension (KIM, 1994). SiC particles present in whisker reinforced alumina
ceramic tools may be responsible for abrasion wear mechanism that generated severe nose
253
wear. It is therefore clear from the machining results that improved properties of the silicon
carbide (SiCw) whisker reinforced alumina ceramic (T10 and T11) tools have negligible
effect on tool performance compared with other commercially available ceramic tools
employed in the machining of titanium alloys.
5.3.5 Nano-grain ceramic tools
Machining with nano-grain ceramic (T12 and T13) tools gave high notch wear rate,
hence lower tool life, relative to other tool grades employed in this study (Figure 4.126). The
Si
3
N
4
base nano-grain B T13 grade gave the better performance in terms of tool wear rate
relative to T12 tool grade when machining Ti-6Al-4V alloy with conventional coolant flow
(Figure 4.127). It has been reported that addition of zirconium oxide (ZrO
2
) improves the
toughness and resistance to fracture of the alumina base tools (BONNEY, 2004). For this
reason T12 tool should theoretically exhibit superior performance than T13 tool because of its
higher edge toughness and hardness (Table 3.3). However, the results shown in Figures 4.126-
4.127 did not confirm this. The relatively poor performance of T12 insert in terms of nose
wear rate (tool life) can be associated with their higher hardness with associated brittleness
compared to Si
3
N
4
base nano-grain B T13 grade (Table 3.3). The higher TiCN content in T12
can also be responsible for accelerated and severe notching due to increased affinity of TiC to
the titanium workpiece material which often leads to catastrophic tool failure at speeds in
excess of 110 m min
-1
(Figure 4.128). Occurrence of catastrophic failure is often influenced
by the stress developed during machining and the toughness property of the tool. Severe
chipping was also observed after machining with tools T12 at cutting speed of 200 m min
-1
, as
illustrated in Figures 4.128. This type of wear occurs on a purely random manner and cannot
be predicted. According to Khamsehzadeh (1991) the high temperatures generated during
machining also promote the development of uneven stress regions on the cutting tool due to
varying microstructure which will increase the stress on the ceramic bond. Alumina ceramic
tools are more susceptible to catastrophic failure due to their low toughness when machining
nickel alloys (KHAMSEHZADEH, 1991). Additionally, it was reported that addition of TiC
resulted to increase in toughness of ceramic tools by increasing their thermal conductivity and
reduce the possibility of the fracture.
254
5.4 Components forces when machining Ti-6Al-4V alloy with different grades of
carbide, PCD, CBN and ceramic tools
The cutting force can be defined as the force exerted by the tool cutting edge on the
workpiece material to promote chip shearing. Cutting force can be resolved into three
components in an orthogonal reference system usually involving the machine tool coordinate
system (Figure 2.18 (a)).
These three components forces are represented by cutting force (F
c
),
feed force (F
f
) and finally the radial force (F
r
) also called the thrust force. Cutting force is
usually the largest force acting on the tool rake face in direction of the cutting velocity. Feed
force acts parallel to the direction of the tool feed. The radial force tends to push the tool away
from the workpiece in a radial direction (DE GARMO; BLACK; KOHSER, 1999).
Understanding of the forces acting on the cutting tool and the study of their behaviour
are of vital importance in the design and manufacture of machine tools and their components
as well as in helping to optimise tool design and controlling the surface finish and surface
integrity of machined components. Additionally, excessive forces can induce vibrations which
are also detrimental to the quality of the machined components hence they are considered as a
machinability index. In this context, more efforts are directed to selecting optimised cutting
conditions to prevent vibration during machining. A curious fact is that cutting forces
generated during machining can either decrease or increase with increasing speed. An
explanation for the first phenomenon is based on the softening of workpiece material as a
result of high temperature generated at the cutting interface (LI; LOW, 1994). As a
consequence, the shear strength of the workpiece material is lowered and hence lower cutting
forces are required at higher speed conditions. On the other hand, cutting forces can also
increase with increasing cutting speed when the wear rate of the tool increases due to the
softening of tool material, hence higher cutting forces are required when machining at high
speed conditions (ZHAN et al., 2000).
Cutting and feed forces can also be influenced by the coolant delivery method,
especially when machining under high pressure coolant supplies. There are, therefore, two
contrasting tendencies on the cutting forces: (i) they tend to decrease them slightly, due to the
marginally shorter sticking zone region produced and (ii) they tend to increase them due to the
shear yield strength as a result of the generation of lower temperatures (MACHADO, 1990).
In the first case, it is well known that the chip-tool contact length decrease with increase in
coolant pressure during machining (PIGOTT; COLWELL (1952), MACHADO (1990),
255
KOVACEVIC; CHERUKUTHOTA; MAZURKIEWICZ (1995), DAHLMAN (2000)). A
hydraulic wedge is created at the tool-workpiece interface that allows the coolant jet to
penetrate the interface deeply with a speed beyond that necessary even for very high speed
machining. This action reduces the tool-chip contact length/area and also changes the chip
flow direction. A resulting force acting upon the chip is created from the pressure in the fluid
wedge. In this case, a shorter sticking region is produced as tool-chip contact is decreased,
hence cutting forces would be expected to decrease. The second tendency is based on the fact
that the yield strength of the work material at the shear zone increases with increase in coolant
pressure. This ability to remove heat when increasing coolant supply compared to
conventional coolant flow was investigated by Machado (1990) who found a significant
reduction of about 175ºC in cutting tool temperature when machining Ti-6Al-4V alloy at a
cutting speed of 30 m min
-1
. In addition to these two tendencies, it has recently been reported
that cutting forces can also be influenced by the generation of reactive forces inherent of
cutting process under high pressure coolant supplies (BONNEY, 2004). This is illustrated in
Figure 5.1. The reactive force is introduced as a result of the coolant jet momentum which
tends to push the tool away from the workpiece material. These reactive forces can generate
severe vibrations during the cutting process, which is detrimental to the integrity of the
machined component. Therefore, it is important to note that, in this study, the reactive forces
were not subtracted from cutting and feed forces recorded values using high pressure coolant
supplies.
0
10
20
30
40
50
0 5 10 15 20
Coolant pressure (MPa)
Reactive component forc
e
(N)
RFx, S = 0 m/min RFz, S = 0 m/min
RFx, S = 110 m/min RFz, S = 110 m/min
Figure 5.1 - Variation of reactive component forces with coolant pressure supply before
machining Ti-6Al-4V alloy.
256
Figures 4.21 shows that cutting forces generally decrease slightly with increase in
coolant pressure when machining with T1 and T4 tool grades whereas the opposite effect is
observed when machining with T2 and T3 tool grades. As often reported, higher cutting
forces generated with T1 and T4 inserts may be attributed to the reduced chip-tool contact
length at high pressure coolant supplies which consequently provided a shorter sticking
region. In case of machining with T2 and T3 tool grades, the relative increase in cutting forces
with coolant supplies may be due to two possible explanations. Firstly, their smaller grain size
relative to T1 and T4 may have reduced their resistance to deformation at higher temperature
(MARI; GONSETH, 1993), thus accelerating tool wear rate and consequently requiring
higher cutting forces. Secondly, the increased ability to remove heat from the cutting zone
with increasing coolant supply, compared to conventional coolant flow may have outweighed
the effect of reducing the contact length, thus leading to increase in cutting forces. Figures
4.21 and 4.22 also show that, in general, cutting and feed forces marginally increased with
cutting speed when machining with all carbides tools tested. Cutting forces recorded with T2
tool grade were generally higher than those recorded with T3 tool in all conditions tested.
This may be attributed to coating of T3 which could have improved lubrication at the cutting
interface, hence requiring lower cutting forces. High cutting and feed forces were generated
when machining with T1 and T4 tool grades in the presence of argon relative to conventional
coolant flow due probably to the poor thermal conductivity of argon which tends to
concentrate more heat at the cutting interface. This accelerates tool wear during machining at
higher cutting speeds due to the softening of tool material, hence higher cutting forces are
required when machining at high speed conditions.
Figures 4.69 and 4.70 show variations in cutting and feed forces, respectively, when
machining Ti-6Al-4V alloy with PCD (T5 and T6) tool grades under various cutting speeds
and coolant pressures. Although PCD tools were employed at higher speed conditions,
comparison of Figures 4.21 and 4.69 show that in general PCD tools generated lower cutting
forces than carbides tools and they vary marginally under all the coolant pressures
investigated. This can be attributed to the higher hardness of PCD tools (Table 3.3), which
tend to retain a sharp and consistent cutting edge with prolong machining at elevated
temperature conditions, therefore requiring lower component forces to perform cutting.
Cutting forces generally increase with increase in cutting speed when machining with T5 and
T6 inserts at lower cutting speeds using conventional coolant supply, contrary to expectation.
257
However, cutting forces generally decrease with increase in coolant pressure when machining
with both T5 and T6 inserts.
Feed forces generated when machining with T5 and T6 inserts are relatively lower than
cutting forces, especially at higher speed conditions, as expected. Feed forces generally
decrease with increase in cutting speed when machining with both T5 and T6 inserts with
conventional coolant flow and increase when machining with high coolant pressures. The
highest feed forces were recorded when machining with T6 insert at the highest coolant
supply of 20.3 MPa (Figure 4.70). Unlike the cutting forces, feed forces increased with
increase in coolant pressure. The higher feed forces recorded at the higher speed conditions
may be attributed to the reactive forces introduced as a result of the high coolant pressure. The
reactive feed forces measured at the various coolant pressures show that they increase with
increasing coolant pressure (Figure 5.1). Figure 5.1 also shows that the reactive forces
increased when the workpiece is rotating. According to Bonney (2004) this phenomenon
occurs because dynamic forces are introduced which increase with increasing cutting speed
for the same mass of workpiece material.
Figures 4.105 and 4.106 show variations in cutting and feed forces, respectively, when
machining Ti-6Al-4V alloy with different CBN (T7,T8,T9) inserts with different coolant
pressures and at various cutting speeds. Lower cutting and feed forces were generated when
machining with conventional coolant flow at a the lowest cutting speed of 150 m min
-1
whereas higher cutting forces were generated when machining with all the CBN grades with
coolant pressure of 11 MPa at a cutting speed of 200 m min
-1
. This may be attributed to the
reactive forces introduced as a result of the coolant pressure especially at higher speed
conditions. In general, lower cutting and feed forces were generated when machining with
CBN tools relative to carbide and PCD tools. Figure 4.105 also shows that relatively lower
cutting and feed forces were also generated when machining with high coolant pressures at
speeds in excess of 200 m min
-1
.
Figures 4.120 and 4.121 show plots of cutting and feed forces, respectively, when
machining Ti-6Al-4V alloy with micron-grain size ceramic (T10 and T11) inserts at different
cutting speeds and under various machining environments. Higher cutting forces were
generated when machining with high coolant pressures compared to argon enriched
environment. This can be attributed to the reactive forces inherent of cutting process with high
coolant pressures (BONNEY, 2004). Additionally, component forces generated during
machining are proportional to stresses at the tool cutting edge. Therefore high compressive
258
forces at the cutting edge will lead to accelerated tool wear and plastic deformation of the tool
edge. These can adversely affect the cutting edge geometry. Prolong machining leads to
accelerated tool wear as a result of high cutting edge temperature (Kear et al., 2001). In
general cutting forces generated with T10 inserts decreased with increasing cutting speed.
Evidence of decreasing cutting forces with increasing cutting speed up to 110 m min
-1
was
reported by Ezugwu et al. (2005) when finish turning Ti-6Al-4V alloy with uncoated carbides
in an argon enriched environment. The increase in cutting forces with increasing cutting speed
in this case may be attributed to the very high wear rate of the ceramic tools (Figures 4.118
and 4.119). The very high wear rate of ceramics is due to their high reactivity with titanium
alloys (LEE (1981), DEARNLEY; GREARSON (1986)) which causes severe notching and
chipping of cutting edge. This tends to increase frictional forces during machining and the
consequent loss/blunting of the sharp cutting edge. This phenomenon also explains the reason
of higher cutting and feed forces generated with ceramics (T10, T11) tool relative to those
generated with carbides, PCD and CBN tools.
Machining Ti-6Al-4V alloy with nano-grain size ceramic (T12 and T13) inserts gave
high component forces (Figure 4.129). Cutting forces generated during machining are higher
than feed forces, as expected. In general component forces increased with increasing in
cutting speed. This may be attributed to high chemical reactivity of ceramic with titanium
alloys which tends to increase frictional forces during machining and the consequent
loss/blunting of the sharp cutting edge. Despite the relative lower component forces generated
relative to other cutting tools, nano-grain size ceramics gave the worst performance,
suggesting that their superplastic flow temperature is low and inadequate for machining
titanium alloys. The very poor performance of nano-ceramic tools when machining Inconel
718 was attributed to this drawback (BONNEY, 2004).
5.5 Surfaces roughness and runout values when machining Ti-6Al-4V alloy with
different grades of carbide, PCD, CBN and ceramic tools
Figures 4.23, 4.71, 4.107, 4.122 and 4.130 show the surface roughness recorded when
machining Ti-6Al-4V alloy with different cutting tools at various cutting conditions. It is
important to note that surface roughness values were recorded after one minute machining
time whereas runout deviation measurements were taken at the end of tool life. Surface
259
roughness values recorded when machining with cemented carbides, PCD and CBN tools
(Figures 4.23, 4.71 and 4.107 respectively) were below the stipulated rejection criterion of 1.6
µm for finishing operation, unlike those recorded with ceramic tools (Figures 4.122 and
4.130). The lowest surface roughness values hence improved surface finish were recorded
when machining with cemented carbides, especially with T1 tool grade, using high coolant
pressures of 11 MPa and 20.3 MPa at cutting speeds up to 120 m min
-1
, while the highest
values were obtained with ceramic T10 tool grade with the various coolant pressures. The
better finishes generated with carbide tools relative to other cutting tools may be attributed to
their larger nose radius of 1.2 mm (Table 3.3). It has been established (KALPAKJIAN;
SCHMID, 2000) that roughness value is closely related to the corner radius and feed rate
given by equation
r
f
R
a
2
0321.0 ×
= (2.14), where f is the feed rate and r is the tool nose
radius (contact radius). According to this equation, the bigger the nose radius the better the
surface finish generated. Therefore, better surface finish is generated when machining Ti-6Al-
4V with carbides with conventional and high pressure coolant supplies. Additionally, the
smooth surface finish generally obtained with carbide inserts when machining under high
pressure coolant supplies can be attributed to the formation of segmented chips. Coolant
supply at high pressures tend to enhance chip segmentation as the chip curl radius is
significantly reduced, hence maximum coolant pressure is restricted only to a smaller area on
the chip (MACHADO (1990), DALHMAN; KAMINSKI (1999), KAMINSKI; DAHLMAN
(1999), DAHLMAN (2000), BONNEY (2004)).
The poor performance of ceramic tools can
be attributed to their high chemical reactivity with titanium alloys which tend to produce
severe abrasive wear, chipping and consequently leading to a loss of the edge sharpness
(Figures 4.118 and 4.119). This adversely affects the surface finish generated during
machining (Figures 4.122 and 4.130). Surface finish deteriorates with increase in cutting
speed when machining with carbides, CBN and ceramic tools (Figures 4.23, 4.107, 4.122 and
4.130, respectively), unlike when machining with both PCD (T5 and T6) inserts which
showed no significant variation in the surface roughness with cutting speed (Figure 4.71). The
relatively low nose wear rates generated when machining with PCD tools suggests that they
maintained their sharp cutting edge for longer periods. It can, however, be seen in Figure 4.71
that there is gradual deterioration of the surface finish with increase in coolant pressure when
machining with larger grain size (T5) insert at cutting speeds up to 230 m min
-1
whereas a
reverse surface roughness pattern was observed when machining with T6 insert. Plots in
260
Figures 4.23, 4.71 and 4.107 show that finish of the machined surfaces are not adversely
affected when machining Ti-6Al-4V with high coolant pressures using carbide, PCD and
CBN inserts under the cutting conditions investigated.
Figures 4.24, 4.72, 4.108, 4.131 show runout deviation values recorded when machining
Ti-6Al-4V alloy with different cutting tools at various cutting speeds and with different
cutting environments. In all cases, the runout values recorded are well below the stipulated
rejection criterion of 100
µm. In general runout increased with increase in cutting speed when
machining with carbides (T1,T2,T3 and T4) and ceramic (T11,T12 and T13) inserts with
conventional coolant flow, unlike PCD (T5 and T6) inserts. The lowest runout values (0.3
µm) were obtained when machining with T4 (carbide) and T5 (PCD) inserts with the highest
coolant pressure of 20.3 MPa. Additionally, minimal variation in recorded runout values were
also obtained when machining with PCD tools with high coolant pressures (Figure 4.72). This
may be attributed to the higher hardness of PCD inserts that ensures more rigidity to the tool-
workpiece system during machining and also due to their lower chemical reactivity with
titanium relative to other cutting tools, hence providing lower nose wear rates. Consequently,
PCD tools can maintain sharp their cutting edge for longer periods, thus ensuring minimal
variation in recorded runout values. However, high runout values were obtained when
machining with CBN and ceramic tools compared to carbide and PCD tools. Ceramic and
CBN tools have high chemical reactivity with titanium alloys and are therefore more
susceptible to accelerated tool wear, especially under higher cutting conditions. Reduction of
hot hardness at elevated temperatures conditions during machining can lead to the weakening
of the inter-particle bond strength and the consequent acceleration of tool wear (NORTH,
1986). Typical tool failure modes observed when machining Ti-6Al-4V alloy with CBN and
ceramic tools are chipping of cutting edge, severe notching and eventually catastrophic tool
failure (Figures 4.97-104 (b), 4.118 (c), (d) and (f)) and 4.119 (b). These types of wear occur
on a purely random manner and cannot be predicted, leading to a loss of the edge sharpness
which adversely affects the dimensional tolerance of a machined component.
5.6 Surface generated of Ti-6Al-4V after machining with carbide and PCD tools
Figures 4.25-4.32, 4.73-4.79 show micrographs of surfaces generated when machining
Ti-6Al-4V alloy with various grades of carbides and PCD tools. In general the surfaces
261
generated with carbide tools consist of well-defined uniform feed marks running
perpendicular to the direction of relative work-tool motion with no evidence of plastic flow.
Irregular feed marks were, however, observed in some cases when machining with the smaller
grain size PCD T6 tool grade at a cutting speed of 175 m min
-1
using conventional coolant
supply and also with a coolant pressure of 11 MPa as illustrated in Figures 4.77 (a) and 4.78
(a), respectively. There is the evidence of localised incipient melting of the machined surfaces
when machining with uncoated carbide T1 tool grade at a higher cutting speed of 130 m min
-1
using conventional coolant flow (Figure 4.25 (b)) and with both grades of PCD tools using
different coolant supplies employed (Figures 4.73 (a), 4.74 (b), 4.76-4.79). Although
distribution of localised incipient melting of machined surface is more severe when
machining with T5 and T6 (PCD) tools, especially when machining with the highest coolant
pressure supply of 20.3 MPa, no evidence of plastic flow was observed. Presence of small
particles (debris) and sometimes surface damages can also be seen on machined surfaces
when machining with carbides and PCD inserts using high coolant pressures (Figures 4.27 (a),
4.28, 4.30, 4.31 (b), 432 (a)-(c), 4.75,4.76 and 4.78). This can be attributed to the smaller and
fragmented chips generated when machining with high coolant pressures which are thrown
out against the workpiece by the high pressure coolant jet. In other cases the surface finish of
machined surfaces are not adversely affected when machining Ti-6A-4V alloy under high
pressure coolant supplies with carbides and PCD tools, as illustrated in Figures 4.23 and 4.71,
for example. Generally surfaces generated under all the finishing conditions investigated
when machining with both PCD inserts were free from other damages such as cracking,
tearing and rupture that are detrimental to machined components. Therefore, surfaces
generated in these trials machined are acceptable and conform to the standard specification
established for machined aerospace components - Rolls-Royce CME 5043.
5.7 Surface hardness after machining Ti-6Al-4V alloy with different grades of carbide,
PCD, CBN and ceramic tools
The plots of microhardness measurements of the machined cross-sections when
machining with carbide (T1,T2,T3,T4) inserts indicate that in general the curves vary
randomly around the hardness range (confidence interval) of the workpiece material and most
of them below the maximum (Max.) Vickers hardness value recorded for the Ti-6Al-4V alloy
262
prior to machining (Figures 4.33-4.42). This suggests that there was no considerable surface
hardening after machining with carbides at the cutting conditions investigated. Surface
hardening up to about 0.4 mm below the machined surfaces of Ti-6Al-4V alloy was observed
after machining with uncoated carbide T1 insert at a cutting speed of 120 m min
-1
using
conventional coolant flow. Evidence of surface hardening up to about 0.25 mm below the
machined surface was also observed when machining with T2 tool using conventional coolant
flow at the lowest speed of 100 m min
-1
. This may be attributed to severe plastic deformation
generated under conventional coolant supply when compressive and shear stresses at the
cutting interface become higher than the yield point of the work material due to high wear
land and cutting temperature (TRENT (1988b), MACHADO; WALLBANK (1994),
DAHLMAN (2000), BONNEY (2004)). Machining with T1 tool using the highest coolant
pressure of 20.3 MPa gave minimum hardness variation with a uniform distribution of
hardness values within the confidence interval of hardness values (Figure 4.36). Analysis of
Figures 4.34-4.37 shows that minimum hardening effect was most accentuated in the
following decreasing order: conventional coolant flow, argon enriched, 11 MPa and 20.3 MPa
coolant pressures respectively. Recorded results clearly show that increasing coolant pressure
generally reduced the hardening effect. Hardening effect in this case is due to high plastic
flow rate combining with the heat generation at the primary shear zone. Efficient coolant
supplies conditions enhance the access of the coolant to the chip-tool interface and contribute
to reducing friction coefficient and the resistance to primary shear zone (SALES (1999),
TRENT; WRIGHT (2000), BONNEY (2004)). As a consequence of this, heat generation is
decreased, resulting to lower temperatures and plastic flow which minimizes the hardening
effect. However, the cutting fluid has negligible access to the tool-workpiece or the tool-chip
interfaces which are under seizure condition when machining at higher speed conditions. The
high temperature generated close to the tool edge during machining generally causes
vaporisation of the cutting fluid (BONNEY, 2004) hence reducing the efficiency of cutting
fluid to remove heat from cutting zone. This may explain the softening of the machined
surface observed when machining with T1 insert at a relatively high speed of 130 m min
-1
using conventional coolant flow. Machining with T3 insert in all cutting conditions
investigated generated hardness values that are uniformly distributed within the confidence
interval with minimum variation of hardness values prior to machining. This, therefore,
suggests that T3 inserts has superior performance in terms of minimum hardness variation
when machining Ti-4Al-4V relative to other carbide inserts employed. Plots of microhardness
263
measurements after machining with coated carbide T4 inserts show evidence of softening of
machined surface when machining in all the cutting conditions investigated (Figures 4.40-
4.42). This softening effect was more pronounced when machining with conventional coolant
flow, as all the hardness values recorded are below the minimum hardness values of the
titanium workpiece prior to machining (Figure 4.40). This softening phenomenon may be
attributed to the over-ageing of titanium alloy at the local surface (CHE-HARON, 2001).
There is longer tool-workpiece contact time when machining at lower cutting speeds. The low
thermal conductivity of titanium alloy encourages retention of the temperature below the
machined surface. Softening of machined surfaces was also observed when machining with
T4 insert under high pressure coolant supplies (Figure 4.41 and 4.42) due probably to the
higher tempering effect generated when machining at higher cutting conditions. Other cause
could be the TiAlN coating of T4 insert. The coating generally acts as a heat barrier, because
it has a much lower thermal conductivity than the tool substrate and the workpiece material.
This tends to reduce heat dissipation rate from the cutting zone and consequently retaining
more heat into the workpiece.
Machining Ti-6Al-4V alloy with PCD T5 and T6 inserts generally gave lower hardness
values beneath the machined surfaces compared to when machining with carbide inserts at the
conditions investigated (Figures 4.80-4.86). It is important to note that cutting speed range
when machining with PCD inserts is 140-250 m min
-1
whereas the range for carbides is 100-
130 m min
-1
. This suggests that machining with PCD tools can lead to softening of the
machined surfaces due to probably to the higher tempering effect caused by the higher
temperature generated at higher speed conditions. This result was unexpected because PCD
inserts with higher thermal conductivity than carbide inserts (Table 3.3) was supposed to
provide higher heat dissipation rate during machining. Increased in cutting speed generally led
to increased hardness when machining with the larger grain size PCD (T5) tool under
conventional coolant flow and under coolant pressures up to 11 MPa (Figures 4.80-4.83), as
expected due to higher temperatures generated at the cutting interface. These Figures also
explains minimum variation in hardness values, with uniform distribution of hardness values
around the confidence interval, obtained when machining with T5 insert at a speed of
175 m min
-1
in all the coolant pressures employed. Machining with higher coolant pressures
generally gave less dispersion of hardness values, i.e. hardness values relatively within the
confidence interval compared to conventional coolant flow (Figures 4.80-4.86). Similar to the
phenomenon that occurred when machining with carbide inserts, here efficient coolant
264
supplies enhance the access of the coolant to the chip-tool interface and contribute to
minimise friction and the resistance to primary shear zone (SALES (1999), TRENT;
WRIGHT (2000), BONNEY (2004)) thereby ensuring improved heat dissipation from cutting
zone.
Machining with CBN (T7, T8 and T9) inserts under various coolant pressures at a
cutting speed of 150 m min
-1
generally gave regular microhardness pattern, i.e. evidence of
softening of machined surface up to about 0.15 mm below the top machined surfaces and a
uniform distribution of hardness values around the minimum and maximum values of the
hardness prior to machining (Figures 4.109-4.111). These values suggest that hardness was
not adversely affected when machining with CBN inserts under the cutting conditions
investigated. The plots also suggest that, in general, hardness depth of the machined surface
decreased with increasing coolant pressure. Again, high pressure coolant technique
contributed to reducing friction coefficient and the resistance to primary shear zone (SALES
(1999), TRENT; WRIGHT (2000), BONNEY (2004)) thereby ensuring improved heat
dissipation from cutting zone. Lower microhardness values were recorded when machining
with T9 grade using all the coolant supplies employed. This may be attributed to heat barrier
property of the TiAlN coating in T9 inserts (Table 3.3) that may have reduced the heat
dissipation rate from cutting zone. This tends to retain more heat in the workpiece leading to
the softening of the machined surface.
Machining with silicon carbide (SiCw) whisker reinforced alumina (T10) ceramic
inserts generally gave hardness values around the minimum and maximum values of the
hardness prior to machining at high coolant pressures while evidence of softening of the
machined surface up to a distance of the about 0.7 mm below the machined surface was
recorded when machining with conventional coolant supply (Figure 4.123). This may be
attributed to two possible hypotheses. The first hypothesis relies on the fact that the high
temperature generated close to the tool edge during machining at conventional coolant flow
generally causes vaporisation of the cutting fluid (BONNEY, 2004) hence reducing the
efficiency of cutting fluid to conduct heat away from cutting zone. The second hypothesis is
based on the excellent cooling capability of the high coolant pressure technique that reduces
temperature at the cutting zone, thereby minimising the quantity of heat could be conducted
into the workpiece.
Additionally, it is important to note that Ti-6Al-4V alloy has two different phases (
α-
and
β phases) which have different hardness values. Therefore, some scattered pattern
265
observed in results of surface hardness values recorded when machining with carbide, PCD,
CBN and ceramic tools at the conditions investigated may be attributed the presence of these
phases.
5.8 Subsurface micrographs after machining Ti-6Al-4V alloy with different grades of
carbide, PCD, CBN and ceramic tools
All the micrographs of the etched machined surfaces generated after machining with
carbide (T1,72,T3,T4) inserts exhibit similar characteristics. The well defined grain
boundaries in Figures 4.43-4.57 clearly show that there were no microstructural alterations
such as plastic deformation, in the subsurface of the machined surfaces. However, a slight
plastic deformation of about 10
µm below the machined surface was observed when
machining with T4 tool grade under high pressure coolant supply of 11 MPa at a cutting
speed of 120 m min
-1
(Figure 4.55 (a)). This may be attributed to the combination of low
thermal conductivity of titanium alloy with the heat retention characteristics of the coating on
T4 insert that can prevent heat developed on a given shear plane from being dissipated. This
leads to softening of work material and causing further shear deformation at the same point
(SHAW, 1984). Astakhov (1999) defined plasticity of a material as its ability to undergo
irreversible plastic deformation when a sufficient external load is applied. The shear
deformation introduced by the cutting action creates a new surface layer with different
structure and properties than the base material. Plastic deformation generally occurs under a
complex non-homogeneous stress, structural changes and a transient temperature field. This is
accompanied by five different energy fluxes: release of elastic energy, crack formation, heat
flow, mass transfer and movement and multiplication of dislocations of material. The change
in properties is due to dislocation of the sub-structure due to residual stresses (induced by the
cutting forces and cutting temperatures) and the consequent strengthening of the material. The
plastic deformation extends to a few microns below the machined surface (BONNEY, 2004).
Machined surfaces produced by conventional machining processes (turning, milling, drilling,
reaming, etc) have a tendency to develop compressive residual stress at depth above 50
µm
from surface. Additionally, this surface status is hardly dependent on the cutting parameters,
tool geometry, tool material and cutting fluid condition. For example, when machining in dry
conditions, the residual stresses generated were always tensile at all depths beneath machined
266
surface whereas when turning with cutting fluids at the same conditions compressive stress
were recorded up to 15
µm beneath the machined surface and below this value, the residual
stress became tensile (ARUNACHALAM; MANNAN; SPOWAGE (2004a),
ARUNACHALAM; MANNAN; SPOWAGE (2004b)). Evidence of plastic deformation at
about 20
µm below the machined surface was reported when machining Inconel 718 with
coated carbide tools under finishing conditions (BONNEY, 2004). A recent study (CHE-
HARON, 2001) reported that a thin layer of disturbed or plastically deformed layer
immediately underneath the machined surface was created after finish turning Ti-6Al-4V
alloy with uncoated carbide inserts under dry conditions This study also reported that
machining with nearly worn or worn tool led to the generation of irregular surface, consisting
of tearing and plastically deformed surface as a result of the high temperature developed at the
chip-tool interface.
Although machining of the titanium alloy with PCD tools suggested softening of
machined surfaces due probably to the high temperature generated when machining at higher
speed conditions, they also generated surfaces of Ti-6Al-4V alloy free from mechanical
damage or microstructural alterations (Figures 4.87-4.93). This suggest that the higher
thermal conductivity of PCD tools relative to other cutting tools employed (Table 3.3) is an
essential tool property for efficient machining of titanium alloys since they can increase heat
dissipation rate from the cutting zone, thus preventing occurrence of plastic deformation
during machining.
The well defined grain boundaries shown in the microstructure of the etched machined
surfaces after machining with CBN (T7, T8 and T9) and ceramic (T10) tools (Figures 4.112-
4.114 and 4.124 (a)-(c)) evidence that there was no microstructural alteration in the
subsurface of machined surfaces after machining at the cutting conditions investigated.
5.9 Chips shapes
Machining with uncoated and coated carbide (T1,T2,T3 and T4) , PCD (T5 and T6),
CBN (T7, T8 and T9) and ceramic (T10 and T11) tools generally produced four different
forms of chips: long tubular chip-shape when machining with T1 and T6 inserts with
conventional coolant flow (Figures 4.58 (a) and 4.94 (e)), snarled chip-shape when machining
with T1, T4, T11 inserts in the presence of argon (Figures 4.58 (b) and (k)) and with T2, T3,
267
T4, T5, T6, T10 inserts under conventional coolant flow (Figures 4.58 (h) and (j), 4.94 (a)
and (e), 4.115 (a), (d), (e), (f) and (g), 4.125 (a), (d)-(e).), partially segmented chip-shape
when machining with carbide (T2 and T3) inserts at high pressure coolant supplies (Figures
4.58 (c) and (g)) and segmented C-shaped chip when machining with T1, T4, T5, T6, T7
inserts at high coolant pressure supplies (Figures 4.58 (d), (e) and (l), 4.94 (b), (c), (d), (f) and
(g))., 4.115 (b) and (c), 4.125 (c)). It has been reported that chip formation process is related
to the cutting tool thermal conductivity that alters tool-chip contact length during machining
(TRENT; WRIGHT, 2000). Continuous and snarled chips are undesirable because they
usually wrap themselves around the workpiece or to get tangled around the tool holder,
adversely affecting the surface finish generated and/or causing tool damage. This can also
hinder the use of advanced manufacturing techniques (unattended machining, for instance) as
well as hindering access of the cutting fluid to the cutting zone and associated disposal
problems that ultimately increase machine tool downtime during production. Long
continuous and, in some cases, snarled chip shapes produced when machining with carbide
tools under conventional coolant flow are in good agreement with the literature (FIELD
(1968), KONIG (1979), TURLEY (1981), TRUCKS (1987), BOOTHROYD; KNIGHT
(1989), MACHADO (1990), MACHADO; WALLBANK (1990)). Chip-tool interface
conditions are the most important factors controlling the rate and the amount of deformation
within the heavily deformed areas that will culminate in the catastrophic thermoplastic shear
process to form the segments. When machining Ti-6Al-4V with T2, T3, T8 and T9 inserts the
high coolant pressure system may have failed to break the chips completely into small
segments. They were partially segmented/ribbon type instead. This suggest that there may be
a critical chip thickness, smaller than the jet geometry created when machining with T1, T2,
T5 and T6 inserts, which is unable to break the swarf. According to Machado (1990), when a
very thin chip is forming it has high flexibility due to its higher elasticity. Thus, the bending
force imposed by the jet is insufficient to break the chip as was obtained with T1 and T4
inserts (Figures 4.58 (c) and (g)).
Chip segmentation was observed when machining at high coolant pressures with T1, T4,
T5, T6, T7 and T10 insert grades. Effective chip segmentation was more visible when
machining with carbides (T1 and T4) and PCD (T5 and T6) inserts. This may be attributed to
the lesser chemical reactivity of titanium with these tools that promote a more stable chip
segmentation process. When a coolant is delivered under high pressure, a hydraulic wedge is
created at the tool-workpiece interface which allows the coolant jet to penetrate the interface
268
deeply with a speed beyond that necessary even for very high speed machining. This tends to
lift up the chip after passing through the deformation zone, thereby reducing the tool-chip
contact length/area as well as changing the chip flow direction. The cantilever effect on the
chip is dependant on the pressure distribution, flow rate and cutting tool. The shape and size
of the pressure distribution enable control of the chip form and flow direction (DALHMAN;
KAMINSKI (1999), KAMINSKI; DAHLMAN (1999), DAHLMAN (2000)). Additionally,
the temperature gradient is reduced by penetration of the high-energy jet into the tool-chip
interface, consequently eliminating the seizure effect (MAZURKIEWICZ; KUBALA;
CHOW, 1989). This action tends to provide adequate lubrication at the tool-chip interface
with a significant reduction in friction (EZUGWU; BONNEY; YAMANE, 2003). These
combined with high velocity coolant flow causes the breakage of the chips into very small
segments.
CHAPTER VI
CONCLUSIONS
1. High Pressure coolant technology has proven that longer tool life can be obtained
when machining difficult-to-machine aerospace alloys (Ti-6Al-4V alloy) with
carbide and PCD inserts, hence improving overall machining productivity
2. PCD (T5 and T6) inserts gave the best performance in terms of tool life when
machining Ti-6Al-4V alloy compared with carbide tools and other cutting tools
employed in all the machining conditions investigated.
3. Coarser grain size carbide (T1 and T4) inserts gave longer tool life than T2 and T3
insert grades in all cutting conditions investigated because the finer grain size tools
have increased solubility of WC in the titanium alloy as the surface area of tool
particles exposed to solution wear increases. The presence of TaC in the
composition of T1 and T4 tools also increase wear resistance when machining the
Ti-6Al-4V alloy.
4. T1 inserts outperformed T4 inserts when machining with conventional coolant flow
and in the presence of argon at all the speed conditions investigated while T4 inserts
generally outperformed T1 inserts when machining with high coolant supply
pressures up to 20.3 MPa.
5. T2 (uncoated) and T3 (coated) carbide tools generally exhibited similar
performance in terms of tool life at the cutting conditions investigated.
6. Coolant pressure has a significant effect on tool wear and hence recorded tool life
when machining Ti-6Al-4V alloy with carbides, PCD and CBN inserts under
finishing conditions.
7. High Pressure coolant technology has proven that longer tool life can be obtained
when machining difficult-to-machine aerospace alloys (Ti-6Al-4V alloy) with
carbide and PCD inserts, hence improving overall machining productivity
270
8. Encouraging tool life can be achieved when machining Ti-6Al-4V alloy with
medium (7 MPa) and high coolant pressures of 11 MPa and 20.3 MPa relative to
conventional coolant flow and in argon enriched environment.
9. On average, tool life increased with increasing coolant pressure for all grades of
carbide inserts employed. The ranking order for carbide tools in terms average gain
in tool life relative to conventional coolant flow is T4, T3, T2 and T1 insert grades.
10. Machining with all carbide tools with 20.3 MPa coolant pressure gave lower tool
life than using a lower coolant pressure of 11 MPa at a cutting speed of
130 m min
-1
. Similar behaviour was observed when machining with PCD tools.
This suggests that 20.3 MPa coolant pressure is above the critical pressure for
machining Ti-6Al-4V alloy under such conditions, therefore optimum cutting
conditions for machining the titanium alloy will depend on the correct combination
of several variables such as coolant pressure, tool grain size and cutting speed.
11. Up to 8 folds improvement in tool life were achieved when machining with PCD
tools relative to carbide inserts using conventional coolant flow.
12. Up to 20 folds improvement in tool life can be achieved when machining with T5
insert under the most aggressive conditions of 230 m min
-1
using high pressure
coolant supply compared with conventional coolant flow.
13. Up to 3 folds and 5 folds improvement in cutting speed can be achieved when
machining with carbide (T1) and PCD (T5) inserts respectively, relative to that
currently achieved in the manufacturing environment.
14. Nose and flank wears are the dominant failure modes when machining Ti-6Al-4V
alloy with all the cutting tools investigated. However, machining with PCD tools
with high coolant pressures can lead to plucking process that can erode the brittle
PCD tool particles.
15. Over 88% reduction in tool wear rate can be achieved when machining with PCD
(T5) tool grade under high pressure coolant supplies relative to conventional
coolant flow at higher cutting speeds in excess of 175 m min
-1
All the grades of
CBN and ceramic inserts gave poor performance when machining Ti-6Al-4V alloy
at all the conditions investigated due to accelerated nose wear and, in some cases,
severe chipping of the cutting edge.
16. Machining with PCD tools generated lower cutting and feed forces than carbides
tools.
271
17. Surface roughness values recorded when machining the Ti-6Al-4V alloy with
carbides and PCD tools are generally below the 1.6 µm rejection criterion for finish
turning, whereas higher values were recorded when machining with CBN and
ceramics inserts.
18. Surfaces generated when machining with all the cutting tools were generally
acceptable and free of physical damages such as tears, laps or cracks in all the
conditions investigated.
19. No significant hardness variation occurred on the machined surfaces when
machining with carbides, PCD and ceramic tools.
20. No evidence of microstructure alteration was observed in the machined subsurface
of Ti-6Al-4V alloy after machining with all the cutting tools employed under the
conditions investigated.
21. Machining with PCD, carbide and CBN tools gave effective chips segmentation
using high pressure coolant technique (discontinuous C-shape chips were
produced).
22. The benefits achieved from these results after employing High Coolant Pressure
Technology in Machining Ti-6Al-4V alloy with carbide and PCD tools in the
production line are:
a) 50% reduction in Cycle times demonstrated on JSF Blisk Turning
b) 5 fold reduction in consumables
c) £ 1.5 million capital saving for JSF Factory and £ 1750.00 saving per machined part
(based on projected cost rate and reduction in cycle time).
CHAPTER VII
RECOMMENDATIONS FOR FURTHER WORK
This research project has shown that a step increase in the machining productivity of a
commercially available titanium-base, Ti-6Al-4V, alloy can be achieved using recently
developed cutting tool materials as well as using high pressure coolant supplies up to
20.3 MPa (203 bar). It was evident that High Coolant Pressure Technique provided a
mechanism of breaking the swarf and outstanding performance when machining the titanium
alloy. This is strongly dependent on a cutting tool (especially carbide and PCD inserts),
cutting speed range, feed rate and depth of cut employed as well as a given high coolant
pressure. Selection of the best combination of cutting tool-speed-coolant pressure has been
successfully implemented in the shop-floor by the collaborating industrial partner, resulting in
a significant reduction in the cost of manufacturing components of jet engines without
compromising their integrity. In this context, determination of an optimum coolant pressure
becomes the major aim. However, this requires an exhaustive number of complete tool life
tests that are time consuming and costly.
From the results obtained in this study, the following future work are recommended:
1. Variation of other cutting parameters such as feed rate and depth of cut would
provide adequate results to establish the optimum coolant pressure (s).
Experimental design techniques have been extensively used to reduce the quantity
of machining tests through application of statistical tools such as ANOVA method,
factorial experiment and Statistic software to determinate the significance of cutting
parameters.
2. The use of these tools would reduce the large number of potentially important
parameters to those that are more significant since it is not economically practical to
perform every possible combination of parameter setting. This would also
273
contribute to developing of a model that can relate primary cutting conditions:
speed, feed rate, depth of cut and coolant pressure.
3. The direction of the coolant jet and nozzle geometry (length, shape, diameter)
investigated in several studies (PIGOTT; COLWELL (1952), SMART; TRENT
(1974), SHAW (1984), MACHADO (1990), KOVACEVIC; CHERUKUTHOTA;
MAZURKIEWICZ (1995), SEAH; LI; LEE (1995), NORIHIKO; AKIO (1998),
DAHLMAN (2000), LÓPEZ DE LACALLE et al. (2000)) was found to have
influence on the tribological conditions present at the tool-workpiece interface, tool
wear, cutting forces when machining various work materials, especially titanium
alloys. This suggests that further tests with different directions of application of
cutting fluid would be valuable.
4. Results from this thesis suggest that different wear mechanisms occurred when
machining with all the cutting tools and cutting conditions employed. X-ray
analysis would be an important technique to enable a better investigation and
analysis of wear mechanisms and possible interactions between elements from the
work material and the cutting tool.
At the time of this investigation few studies were reported about the use of high
pressure cooling technology in milling of titanium alloys. Therefore, it will be wise for
future research work to exploit this technology in mill
CHAPTER VIII
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APPENDIX
LIST OF PUBLICATIONS FROM THIS STUDY
Refereed Journals
1. E.O. Ezugwu, J. Bonney, R.B. Da Silva and Á.R. Machado, Evaluation of the
Performance of Different Nano-Ceramic Tool Grades when Machining Nickel-Base,
Inconel 718, Alloy, Journal of the Brazilian Society of Mechanical Science & Engineering
– ABCM / RBCM. Vol. XXVI, No. 1, January-March 2004, pp. 12-16.
2. E.O. Ezugwu, R.B. Da Silva, J. Bonney and Á.R. Machado, The Effect of Argon Enriched
Environment in High Speed Machining of Titanium Alloy, Tribology Transactions –
Society of Tribologists and Lubrication Engineers (STLE) – Vol. 48: pp. 18-23, 2005,
ISBN: 0569-8197.
3. E.O. Ezugwu, R.B. Da Silva, J. Bonney and Á.R. Machado, Evaluation of the
Performance of CBN tools when turning Ti-6Al-4V alloy with high pressure coolant
supplies, International Journal of Machine Tool and Manufacture, Vol. 45, July 2005,
pp.1009-1014.
4. E.O. Ezugwu, D.A. Fadare, J. Bonney, R.B. Da Silva and W. F. Sales, Modelling of The
Correlation between Cutting and Process Parameters in High Speed Machining of Inconel
718 Alloy using Artificial Neural Network, International Journal of Machine Tool and
Manufacture, Vol. 45, 2005, pp. 1375-1385, ISBN: 0890-6955.
299
Refereed Conferences
1. E.O. Ezugwu, J. Bonney, R.B. Da Silva and Á.R. Machado, Evaluation of the
Performance of Different Nano-Ceramic Tool Grades when Machining Nickel-Base,
Inconel 718, Alloy, Presented at the Second Brazilian Manufacturing Engineering
Conference (II COBEF) in Uberlândia-MG, Brazil, May 2003, CD-Rom.
2. E.O. Ezugwu, R.B. Da Silva, J. Bonney and Á.R. Machado, The Effect of Argon Enriched
Environment in High Speed Machining of Titanium Alloy, Presented at the Society of
Tribologists and Lubrication Engineers, 59th. Annual Meeting & Exhibition, in Toronto,
Canada, 17-20 May 2004.
3. E.O. Ezugwu, D.A. Fadare, R.B. Da Silva, J. Bonney and A. Shabazz Nelson, Wear
prediction of Uncoated Carbide Tool during High Speed Turning of Ti-6Al-4V alloy
using Artificial Neural Network, Proceedings of Second International Conference on
Manufacturing Research – ICMR in Sheffield Hallam University, England, pp. 36-41, 7-9
September 2004.
4. E.O. Ezugwu, D.A. Fadare, R.B. Da Silva and J. Bonney, Application of Artificial Neural
Networks for Tool Condition Monitoring when Turning Ti-6Al-4V Alloy with High
Pressure Coolant Supply, Presented at the 7th. International Conference on Progress of
Machining Technology in Zuzhou, China 8-11 December 2004.
5. R.B. Da Silva, E.O. Ezugwu, J. Bonney, Á.R. Machado and A.M. Reis, High Speed
Turning of Ti-6Al-4V alloy with Coated Carbides tools in an Argon Enriched
Environment, Presented at the 38th. CIRP – International Seminar on Manufacturing
System in Florianópolis–SC, Brazil, 16-18 May 2005.
6. R.B. Da Silva, Á.R. Machado, E.O. Ezugwu and J. Bonney, Evaluation of the
Machinability of Ti-6Al-4V alloy with (SiCw) Whisker Reinforced Alumina Ceramic
Cutting Tool under Various Cooling Environments, Presented at the 18th. International
Congress of Mechanical Engineering (COBEM 2005) in Ouro Preto–MG, Brazil, 06-11
November 2005.
7. R.B. Da Silva, Á.R. Machado, E.O. Ezugwu and J. Bonney, Increasing productivity when
machining Ti-6Al-4V alloy in HSM under high pressure coolant supply, presented at the
Fifth International Conference on High Speed Machining (HSM) in Metz, France,13th.-
15th March 2005
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