(RWTHedition) Fritz Klocke (auth )-Manufacturing Processes 1_ Cutting-Springer-Verlag Berlin Heidelberg (2011)

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RWTH edition
RWTH Aachen
Fritz Klocke
Manufacturing Processes 1
Cutting
Translated by Aaron Kuchle
123
Professor Dr.-Ing. Dr.-Ing. E.h. Dr. h.c. Dr. h.c. Fritz Klocke
Laboratory for Machine Tools and Production Engineering
RWTH Aachen University
Steinbachstr. 19
52074 Aachen
Germany
f.klocke@wzl.rwth-aachen.de
Translated by
Aaron Kuchle
Yeungnam University
Korean Language Institute
214-1 Dae-dong Gyeongsan Gyeongbuk 712-749
Korea
aaronkuchle@yahoo.com
ISSN 1865-0899 e-ISSN 1865-0902
ISBN 978-3-642-11978-1 e-ISBN 978-3-642-11979-8
DOI 10.1007/978-3-642-11979-8
Springer Heidelberg Dordrecht London New York
Library of Congress Control Number: 2011925556
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Springer is part of Springer Science+Business Media (www.springer.com)
Preamble About Compendium “Manufacturing
Processes”
Key factors for quality and economic efficiency of industrial production are the
choice of the manufacturing processes and their design. Manufacturing Technology
is an elemental part of the fundamental knowledge of machining engineers. Also
design engineers have to gain knowledge in this field, since they have high respon-
sibility for the manufacturing costs. However, the students as well as practising
experts who are willing to enhance their knowledge have the problem to collect
information. To the current day there is no extensive, but still clear description of
manufacturing processes focussing on the technology itself.
In order to counter this necessity the compendium at hand is supposed to present
an overall picture of the most common machining and non-machining manufac-
turing processes. Additional to the description of the techniques these volumes are
desired to deliver an insight in the underlying physical principles whenever it is
necessary for the understanding of the processes.
The apportionment of the compendium “Manufacturing Processes” into
Volume 1: Cutting
Volume 2: Grinding, Honing, Lapping
Volume 3: Electrical Erosion and Hybride Processes
Volume 4: Forming
Volume 5: Casting, Sintering, Rapid Prototyping
groups techniques with similar active principles together.
In front of the first volume is placed a technique-spanning section to the toler-
ances and questions of the workpiece measuring techniques used in manufacturing.
Within the individual volumes was tried to avoid an encyclopaedic listing of the
techniques. The book series are primarily intended for junior scientists in the fields
of manufacturing technology and construction. In addition, the practitioner will be
able to refurbish or extend his knowledge. The variety of manufacturing problems
v
vi Preamble About Compendium “Manufacturing Processes”
is as large as the multiplicity of the products, and alone with text book wise sayings
manufacturing questions are not to be solved. We wish that this book offers starting
points and ways to its readers, on which they can come up with successful solutions
by engineering thinking.
Aachen, Germany Fritz Klocke
September 2010
Preamble for Volume 1 “Cutting”
The available volume treats machining with geometrically defined cutting edges.
It approaches both the practical engineer and the student of engineering sciences.
This book is based on the lecture “Manufacturing Technologies I and II” and the
pertinent exercises, which are held at RWTH Aachen University. The arrangement
of the book results from the experiences, which were gained when lecturing in the
chronology of the course.
The structure of the book is oriented to a large extent at didactical criteria. The
first section deals with the metrology and the workpiece quality, followed by a chap-
ter about the fundamentals of cutting. Based on the common technological principle
of the cutting processes and their variants, the second section deals with the mech-
anisms occurring at the cutting edge and the resulting loads on the tools. The next
section deals with the demands on the cutting materials as well as their manufac-
turing processes and application. The text presents topics like the use of the finite
element method for cutting processes as well as cutting fluids. A detailed chapter
about the machinability of the most important work piece materials, such as steel
and cast iron, light metal alloy and high temperature resistant materials imparts a
broad knowledge of the alternating coactions of work piece material, the cutting
material and the process parameters.
The necessity for process control, used sensor systems and application possi-
bilities are regarded. Due to the increasing relevance and application in industrial
manufacturing plants, monitoring systems are discussed in their own chapter.
Finally, the different cutting processes and their kinematic characteristics are
presented in detail.
For their cooperation with the compilation of the available edition I would
like to thank my co-workers, Dr.-Ing. M. Abouridouane, Dr.-Ing. K. Gerschwiler,
Dipl.-Ing. M. Arft, Dipl.-Ing. C. Essig, Dipl.-Ing. P. Frank, Dipl.-Ing. A. Krämer,
Dipl.-Ing. D. Lung, Dipl.-Ing. H. Sangermann, Dipl.-Ing. R. Schlosser, Dipl.-Ing.
vii
viii Preamble for Volume 1 “Cutting”
P. Vogtel, as well as Dipl.-Ing. S. E. Cordes, who was additionally responsible for
the coordination of the work on this book.
Further I would like to thank the many former co-workers, who participated in
the past German editions and now have leading positions in industry and research
companies.
Aachen, Germany Fritz Klocke
September 2010
Contents
1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1
2 Metrology and Workpiece Quality . . . . . . . . . . . . . . . . . . 3
2.1 Manufacturing Disturbances and Manufacturing History . . . . 3
2.2 Measuring and Testing . . . . . . . . . . . . . . . . . . . . . . 8
2.2.1 Measurement Errors . . . . . . . . . . . . . . . . . . 9
2.2.2 Macro- and Microgeometry of Components . . . . . . 10
2.3 Length Testing Devices . . . . . . . . . . . . . . . . . . . . . . 13
2.3.1 Material Measures . . . . . . . . . . . . . . . . . . . 13
2.3.2 Gauges . . . . . . . . . . . . . . . . . . . . . . . . . 13
2.3.3 Indicating Measuring Instruments . . . . . . . . . . . 14
2.4 Surface Inspection . . . . . . . . . . . . . . . . . . . . . . . . 21
2.4.1 Surface Parameters . . . . . . . . . . . . . . . . . . . 22
2.4.2 Subjective Surface Inspection . . . . . . . . . . . . . 26
2.4.3 Surface Measurement . . . . . . . . . . . . . . . . . . 27
2.5 Inspection of the Workpiece Rim . . . . . . . . . . . . . . . . 30
2.5.1 Surface Layers . . . . . . . . . . . . . . . . . . . . . 31
2.5.2 Inspection of the Surface Rim Zone . . . . . . . . . . 33
3 Fundamentals of Cutting . . . . . . . . . . . . . . . . . . . . . . . . 39
3.1 The Cutting Part – Concepts and Terms . . . . . . . . . . . . . 39
3.2 Reference Systems . . . . . . . . . . . . . . . . . . . . . . . . 42
3.3 Basic Process Variants. . . . . . . . . . . . . . . . . . . . . . 45
3.4 Chip Formation . . . . . . . . . . . . . . . . . . . . . . . . . . 47
3.4.1 The Cutting Process . . . . . . . . . . . . . . . . . . 47
3.4.2 Different Types of Chip Formation . . . . . . . . . . . 49
3.5 Kinematic Surface Roughness . . . . . . . . . . . . . . . . . . 54
3.6 Mechanical and Thermal Strain on the Cutting Section . . . . . 55
3.6.1 Influence of the Geometry of the Cutting Section . . . 68
3.7 Wear . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 71
3.7.1 Wear Mechanisms . . . . . . . . . . . . . . . . . . . 71
3.7.2 Causes of Wear . . . . . . . . . . . . . . . . . . . . . 74
3.7.3 Forms and Dimensions of Wear . . . . . . . . . . . . 85
ix
x Contents
3.8 Cutting Theory . . . . . . . . . . . . . . . . . . . . . . . . . . 87
3.8.1 Shear Plane Theory . . . . . . . . . . . . . . . . . . . 87
3.8.2 Application of Plasticity Theory to Cutting . . . . . . 92
4 Cutting Tool Materials and Tools . . . . . . . . . . . . . . . . . . . 95
4.1 Overview of Cutting Tool Materials . . . . . . . . . . . . . . . 95
4.1.1 Classification of Hard Cutting Tool Materials . . . . . 97
4.2 Tool Steels . . . . . . . . . . . . . . . . . . . . . . . . . . . . 99
4.2.1 Cold Work Steels . . . . . . . . . . . . . . . . . . . . 100
4.2.2 High Speed Steels . . . . . . . . . . . . . . . . . . . 101
4.3 Cemented Carbides . . . . . . . . . . . . . . . . . . . . . . . . 110
4.3.1 Historical Development . . . . . . . . . . . . . . . . 111
4.3.2 Cemented Carbide Production . . . . . . . . . . . . . 112
4.3.3 Components of Cemented Carbides and Their
Properties . . . . . . . . . . . . . . . . . . . . . . . . 114
4.3.4 Microstructure . . . . . . . . . . . . . . . . . . . . . 116
4.3.5 Classification of Cemented Carbides . . . . . . . . . . 117
4.3.6 Uncoated Cemented Carbides and Cermets . . . . . . 123
4.3.7 Function Gradient Cemented Carbides . . . . . . . . . 123
4.3.8 Coated Cemented Carbides and Cermets . . . . . . . . 126
4.4 Coatings . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 127
4.4.1 Coating Methods . . . . . . . . . . . . . . . . . . . . 128
4.4.2 Specific Properties of Hard Material Coatings . . . . . 149
4.4.3 Substrate Pretreatment . . . . . . . . . . . . . . . . . 161
4.5 Ceramic Cutting Tool Materials . . . . . . . . . . . . . . . . . 161
4.5.1 Cutting Ceramics . . . . . . . . . . . . . . . . . . . . 162
4.5.2 Superhard Non-metallic Cutting Tool Materials . . . . 173
4.6 Tool Designs . . . . . . . . . . . . . . . . . . . . . . . . . . . 184
4.6.1 Solid Tools . . . . . . . . . . . . . . . . . . . . . . . 185
4.6.2 Tools with Inserts . . . . . . . . . . . . . . . . . . . . 187
4.7 Tool Preparation . . . . . . . . . . . . . . . . . . . . . . . . . 193
5 Finite Element Method (FEM) . . . . . . . . . . . . . . . . . . . . 197
5.1 Basic Concepts of FEM . . . . . . . . . . . . . . . . . . . . . 197
5.2 Lagrangian and Eulerian Considerations of the Continuum . . . 199
5.3 Explicit and Implicit Methods of Solution . . . . . . . . . . . . 199
5.4 Combined Thermal and Mechanical FEA . . . . . . . . . . . . 200
5.5 Nonlinearities . . . . . . . . . . . . . . . . . . . . . . . . . . . 200
5.6 Material Laws . . . . . . . . . . . . . . . . . . . . . . . . . . 201
5.7 Software . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 204
5.8 Hardware . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 205
5.9 Phases of a Finite Element Analysis (FEA) . . . . . . . . . . . 205
5.10 The Use of FEM in Cutting Technology . . . . . . . . . . . . . 206
5.10.1 Continuous Chip Simulation . . . . . . . . . . . . . . 207
5.10.2 Segmented Chip Simulation . . . . . . . . . . . . . . 209
5.10.3 Simulation of the Cutting Process . . . . . . . . . . . 212
Contents xi
6 Cutting Fluids . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 219
6.1 The Functions of Cutting Fluids . . . . . . . . . . . . . . . . . 219
6.2 Types of Cutting Fluids . . . . . . . . . . . . . . . . . . . . . . 219
6.2.1 Non Water-Miscible Cutting Fluids . . . . . . . . . . 220
6.2.2 Water-Based Cutting Fluids . . . . . . . . . . . . . . 221
6.3 Guidelines on the Use of Coolant Emulsions . . . . . . . . . . 222
6.4 Effects of the Cutting Fluid on the Machining Process . . . . . 223
6.5 Selection of Cutting Fluids . . . . . . . . . . . . . . . . . . . . 226
6.6 Reducing or Avoiding the Use of Cutting Fluids . . . . . . . . . 227
6.6.1 Reducing Cutting Fluids . . . . . . . . . . . . . . . . 228
6.6.2 Minimum Quantity Cooling Lubrication (MQCL) . . . 229
6.6.3 Avoiding Cutting Fluids . . . . . . . . . . . . . . . . 233
7 Tool Life Behaviour . . . . . . . . . . . . . . . . . . . . . . . . . . . 237
7.1 Determining Tool Life Parameters . . . . . . . . . . . . . . . . 239
7.2 Machinability . . . . . . . . . . . . . . . . . . . . . . . . . . . 239
7.2.1 Tool Life . . . . . . . . . . . . . . . . . . . . . . . . 240
7.2.2 Resultant Force . . . . . . . . . . . . . . . . . . . . . 243
7.2.3 Surface Quality . . . . . . . . . . . . . . . . . . . . . 245
7.2.4 Chip Form . . . . . . . . . . . . . . . . . . . . . . . 248
7.2.5 Cutting Speed . . . . . . . . . . . . . . . . . . . . . . 248
7.3 The Machinability of Steel Materials . . . . . . . . . . . . . . 249
7.3.1 Metallographic Constituents . . . . . . . . . . . . . . 250
7.3.2 Carbon Content . . . . . . . . . . . . . . . . . . . . . 252
7.3.3 Alloying Elements . . . . . . . . . . . . . . . . . . . 254
7.3.4 Types of Heat Treatment . . . . . . . . . . . . . . . . 256
7.4 Machinability of Various Steel Materials . . . . . . . . . . . . 263
7.4.1 Machining Steels . . . . . . . . . . . . . . . . . . . . 263
7.4.2 Case-Hardened Steels . . . . . . . . . . . . . . . . . 266
7.4.3 Heat-Treatable Steels . . . . . . . . . . . . . . . . . . 267
7.4.4 Nitrided Steels . . . . . . . . . . . . . . . . . . . . . 268
7.4.5 Tool Steels . . . . . . . . . . . . . . . . . . . . . . . 269
7.4.6 Hardened Steels . . . . . . . . . . . . . . . . . . . . 271
7.4.7 Non-rusting Steels . . . . . . . . . . . . . . . . . . . 274
7.5 Machinability of Cast Iron . . . . . . . . . . . . . . . . . . . . 280
7.5.1 White Cast Iron . . . . . . . . . . . . . . . . . . . . . 280
7.5.2 Grey Cast Iron . . . . . . . . . . . . . . . . . . . . . 283
7.6 Machinability of Non-ferrous Metals . . . . . . . . . . . . . . 290
7.6.1 Aluminium Alloys . . . . . . . . . . . . . . . . . . . 290
7.6.2 Magnesium Alloys . . . . . . . . . . . . . . . . . . . 293
7.6.3 Titanium Alloys . . . . . . . . . . . . . . . . . . . . 296
7.6.4 Copper Alloys . . . . . . . . . . . . . . . . . . . . . 311
7.6.5 Nickel Alloys . . . . . . . . . . . . . . . . . . . . . . 314
7.7 Machinability of Non-metals . . . . . . . . . . . . . . . . . . . 327
7.7.1 Graphite . . . . . . . . . . . . . . . . . . . . . . . . . 327
7.7.2 Fibre-Reinforced Plastics . . . . . . . . . . . . . . . . 331
xii Contents
8 Process Design and Process Monitoring . . . . . . . . . . . . . . . 339
8.1 Finding Economical Cutting Parameters . . . . . . . . . . . . . 339
8.1.1 Cutting Parameter Limits . . . . . . . . . . . . . . . . 339
8.1.2 Optimizing the Cutting Parameters . . . . . . . . . . . 342
8.1.3 Calculating the Machine Hour-Rate . . . . . . . . . . 350
8.1.4 Planning Methods and Tools . . . . . . . . . . . . . . 351
8.2 Process Monitoring . . . . . . . . . . . . . . . . . . . . . . . . 352
8.2.1 Sensors for Process Monitoring . . . . . . . . . . . . 354
8.2.2 Signal Processing and Monitoring Strategies . . . . . 371
9 Processes with Rotational Primary Movement . . . . . . . . . . . . 383
9.1 Turning . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 383
9.1.1 Face Turning . . . . . . . . . . . . . . . . . . . . . . 385
9.1.2 Cylindrical Turning . . . . . . . . . . . . . . . . . . . 386
9.1.3 Helical Turning . . . . . . . . . . . . . .. . . . . . . 387
9.1.4 Profile Turning . . . . . . . . . . . . . . . . . . . . . 388
9.1.5 Form Turning . . . . . . . . . . . . . . . . . . . . . . 389
9.1.6 Further Process Variants . . . . . . . . . . . . . . . . 390
9.2 Milling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 390
9.2.1 Process Variants, Specific Characteristics and Tools . . 395
9.3 Drilling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 416
9.3.1 Profile Counterboring . . . . . . . . . . . . . . . . . . 416
9.3.2 Rotary Drilling . . . . . . . . . . . . . . . . . . . . . 418
9.4 Sawing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 439
9.4.1 Bandsawing . . . . . . . . . . . . . . . . . . . . . . . 439
9.4.2 Hacksawing . . . . . . . . . . . . . . . . . . . . . . . 442
9.4.3 Circular Sawing . . . . . . . . . . . . . . . . . . . . 442
10 Processes with Translatory Primary Movement . . . . . . . . . . . 445
10.1 Broaching . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 445
10.1.1 Face and Circular Broaching . . . . . . . . . . . . . . 446
10.1.2 Profile Broaching . . . . . . . . . . . . . . . . . . . . 449
10.1.3 Form Broaching . . . . . . . . . . . . . . . . . . . . 451
10.2 Shaving . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 454
10.3 Planing and Shaping . . . . . . . . . . . . . . . . . . . . . . . 457
10.3.1 Face Shaping and Face Planing . . . . . . . . . . . . 458
10.3.2 Gear Shaping . . . . . . . . . . . . . . . . . . . . . . 459
10.3.3 Gear Planing . . . . . . . . . . . . . . . . . . . . . . 466
References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 469
Subject Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 497
Symbols and Abbreviations
Capital Letters
A mm2 Cross section
A % Stretch ratio
Ac1 Transformation temperature (Start of austenite formation)
Ac3 Transformation temperature (End of austenite formation)
Aα mm2 Flank face
A′α mm2 Minor flank face
Aγ mm2 Rake face
A� mm2 Shear face
AFK Aramid Faser reinforced plastics
AIP Arc Ion Plating
AV mm Offset
BAZ Machining center
Cf min Axis intercept of the expanded TAYLOR equation,
f = 1 mm
Cv min Axis intercept of the expanded TAYLOR equation,
vc = 1 m/min
CBN Cubic boron nitride
CFK Carbon fibre reinforced plastics (CFRP)
CNC Computerised Numerical Control
CVD Chemical Vapour Deposition
CWT Continuous wavelet transform
D mm Tool diameter
DC Direct current
DLC Diamond-like carbon
DMS Strain gauge
DP Polycrystalline diamond
DWT Discrete wavelet transform
E Young’s modulus
EB-PVD Electron beam-PVD
xiii
xiv Symbols and Abbreviations
ECM Electro chemical machining
EDX Energy-dispersive X-ray spectroscopy
F N Resultant force
Fa N Active force
Fc N Cutting force
FD N Thrust force
Ff N Feed force
Fp N Passive force
F� N Shear force
FFT Fast Fourier transform
FVK Fibre-reinforced plastic (FRP)
GFK Glass-reinforced plastic (GRP)
H mm Stroke length
HB Brinell hardness
HC Diamond coated cemented carbide
HF Fine grain cemented carbide
HF-CVD Hot-Filament-Process-CVD
HM Cemented carbide
HRC Rockwell hardness
HSC High Speed Cutting
HSS High Speed Steel
HT-CVD High temperature-CVD
HV Vicker hardness
KA C Imputed amortizations
KE C Energy costs
KI C Maintenance costs
KF,min C/St. Minimal manufacturing costs per workpiece
KMH C/h Machine hour-rate
KML C Machine and labour cost hourly rate
KR C Room costs
KZ C Imputed interest
KB mm Crater width
KM mm Crater center distance
KSS Cutting fluid
KT mm Crater depth
L C/h Hourly wage
LARC Lateral Rotating Arc
MQCL Quantity Cooling Lubrication
MQC Minimum quantity cooling
MQS Minimum quantity lubrication
MQL Minimum quantity lubrication
MSIP Magnetron Sputter Ion Plating
MT-CVD Moderate Temperature-CVD
NC Numerical control
Symbols and Abbreviations xv
NE-Metalle Non-ferrous metals
PA-CVD Plasmaactivated-CVD
Pc W Cutting power
Pe W Effective power
Pf Assumed working plane
Pf W Feed power
Pn Tool cutting edge normal plane
Pne Effective cutting edge normal plane
Po Tool orthogonal plane
Pp Tool back plane
Pr Tool reference plane
Ps Cutting edge plane
PW W Effective power
PKD Polycrystalline diamond
PM Powder metallurgy
PM-HSS Powder Metallurgy High Speed Steel
PVD Physical Vapour Deposition
Q C Charge
R � Electrical resistance
Ra μm Average roughness
Re N/mm2 Elastic limit
REM Scanning electron microscope (SEM)
Rkin μm Surface quality
Rm N/mm2 Tensile strength
Rp0,2 N/mm2 0.2%–yield strength
Rt μm (Depth of) Surface roughness
Rz μm Averaged surface roughness
RF Radio Frequency
S Major cutting edge
S′ Minor cutting edge
SKV μm Rounded cutting edge
STFT Short-Time Fourier Analysis
SVα mm Displacement of the cutting edge in the direction of
the flank face
SVγ mm Displacement of the cutting edge in the direction of
the rake face
T min Tool life
T ◦C Temperature
Tm ◦C Melting temperature
Tr ◦C Room temperature
Tsp ◦C Temperature on the chip upper side
U V Voltage
U0 V Feeding voltage
VB mm Width of flank wear land
xvi Symbols and Abbreviations
VBmax mm Maximal width of flank wear land
VHM Cemented Carbide
Wc J Cutting work
We J Effective work
Wf J Feed work
WZ-D mm Tool diameter
WSP Insert
Z Atomic number
ZTU Time Temperature Transformation (TTT)
Small Letters
ae mm Width of cut
ap mm Depth of cut
apmax mm Maximum depth of cut
apok mm Optimal-cost depth of cut
a-C Hydrogenfree amorphous carbon film
a-C:H Hydrogenated amorphous carbon film
at.-% Atomic ratio
B mm Width of undeformed chip
bfα mm Width of chamfer
C m/min Speed of sound
cp Process capability index
cpk Rectified process capability index
dij pC/N Piezoelectric coefficient
dW/dt m/min Growth of wear
f mm Feed
fa mm Axial feed
fok mm Optimal-cost feed
fmax mm Maximal feed
fmin mm Minimal feed
fr mm Radial feed
fz mm Feed per tooth/cutting edge
f0 1/s Eigenfrequency
H mm Undeformed chip thickness
hch mm Chip thickness
hcu,max μm Maximal chip thickness
hmin mm Minimal chip thickness
K Strain sensitivity
kc1.1 N/mm2 Specific cutting force, b = h = 1 mm
kf1.1 N/mm2 Specific feed force, b = h = 1 mm
ki N/mm2 Specific resultant force
kp1.1 N/mm2 Specific passive force, b = h = 1 mm
Symbols and Abbreviations xvii
lc m Cutting path
lf m Feed path
mc Exponent of the specific cutting force
mn mm Normal module
pKSS bar Cutting fluid pressure
rε mm Corner radius
rß mm Rounded cutting edge radius
sd Coating thickness
T mm Pitch
T s Time
te,min min Minimal time per unit
tg min Basic time
tv min Additional time
te min Time per unit
ter min Recovery time
th min Main process time
tn min Auxiliary process time
tr min Setting-up time
ta-C Tetrahedral hydrogenated amorphous carbon film
vc m/min Cutting velocity
vch m/min Chip velocity
vf m/min Feed velocity
vfax mm/min Axiale feed velocity
ve m/min Effective velocity
x mm Profile offset
z Number of teeth
Greek Letters
α ◦ Tool clearance angle
αeff ◦ Effective tool clearance angle
αf ◦ First tool orthogonal clearance (Broaching)
αF
◦ Clearance angle at the pitch diameter (Gear shaping)
αo
◦ Tool clearance angle
αoe
◦ Working clearance angle
β ◦ Wedge angle
βo
◦ Wedge angle
βoe
◦ Working wedge angle
γ o
◦ Tool orthogonal rake angle
γ n
◦ Tool normal rake angle
γ ne
◦ Working rake angle
γ oe
◦ Wirk-Orthogonalspanwinkel
γ eff ◦ Effective rake angle
xviii Symbols and Abbreviations
ε Strain
ε¯ Shear strain
˙¯ε s–1 Shear strain speed
ε˙ 1/s Strain rate
εk Critical strain
εr
◦ Tool included angle
η ◦ Effective cutting speed angle
ϑ ◦C Temperature
κ r
◦ Tool cutting edge angle
κ re
◦ Working cutting edge angle
λs
◦ Tool cutting edge inclination
λse
◦ Working cutting edge inclination angle
μ Coefficient of friction
φ ◦ Phase
φ ◦ Shear angle
φf ◦ Slide angle
ϕ ◦ Feed motion angle
ρ kg/m3 Density
σN/mm2 Stress
σn N/mm2 Normal stress
ω rad/s Angular speed
Chapter 1
Introduction
From about 12,000 to 50,000 years ago, human beings could already adapt stone
tools with intentionally produced edges to specific machining tasks by varying the
geometry of the cutting edge, as is shown in early tool findings from the Palaeolithic
Age (Fig. 1.1).
The discovery of how to extract metals such as copper, tin and iron was a huge
milestone in the history of material and manufacturing technology. Since about
700 B.C., tools were almost exclusively made of iron. At the beginning of the 17th
century, constant improvements in iron smelting led to the preferred use of iron and
steel as construction materials instead of other metals known then and in place of
wood, which had predominated until then.
Inspired by the growth of the textile industry and the discovery of the steam
engine, there was increased exploration into manufacturing technology at the begin-
ning of the 19th century, leading by the second half of the 19th century to the
first systematic investigations into cutting methods and initiating a completely new
research area. At the end of the 19th century further research led to the discovery
of new cutting tool materials and, at the turn of the century, to the development of
high speed steel by F.W. TAYLOR, a significant contribution to the history of manu-
facturing technology [Tayl07]. In light of these developments, SCHLESINGER said:
“The dividends may lie on the cutting edge of steel, but the speed of these cutting
edges is a function of the machine moving them, so as wages increase, the cutting
machine is a trump card.” [Schl11].
Subsequent research followed in this direction, leading to the development of
cemented carbide in 1923 by SCHRÖTER [Schr23, DRP25] at the SOCIETY FOR
ELECTRIC LIGHTING as well as its use in machining by FRIEDR. KRUPP AG as
WIDIA cemented carbide. Following this was the invention of oxide-ceramic cutting
tool materials and their application in cutting after 1938 by OSENBERG as well
as WENTORF’s development and synthesis of the superhard cutting tool material
cBN (cubic-crystalline boron nitride) in 1956 [Went57]. The refinement of cutting
tools with wear-resistant cemented carbide coats starting in 1968 by the companies
Sandvik Coromant and Friedr. Krupp AG was a major contribution to improving
productivity and economy [Sche88].
In order to fulfill the constantly increasing requirements made on workpiece
quality and to make machining processes more economical, all the influencing
1F. Klocke, Manufacturing Processes 1, RWTH edition,
DOI 10.1007/978-3-642-11979-8_1, C© Springer-Verlag Berlin Heidelberg 2011
2 1 Introduction
Saw
Drill
End scraper
Gravettian point Notched point
Stemmed-base
hafted scraper
Side scraperRound scraperDouble tines
Fig. 1.1 Flint tools, acc. to KERND’L [Kern72]
parameters relevant for productivity and workpiece quality must be considered:
cutting edge geometry, process kinematics, cutting parameters, materials, cutting
tool materials and their coatings, wear behaviour as well as auxiliary materials. In
this pursuit, the following quotation by SCHLESINGER can act as a guideline: “The
machine tool must be designed with a view to the place it is used”.
As an introduction to the subject matter, the compendium will be preceded by a
general chapter providing a survey of the most important attributes of workpieces
and their test methods.
Chapter 2
Metrology and Workpiece Quality
2.1 Manufacturing Disturbances and Manufacturing History
The basic task of manufacturing is to provide workpieces with specified quality
characteristics in the required quantity in the most time- and cost-efficient way
possible. Every manufacturing process is affected by variable disturbances, which
can be both external and internal (occurring within the process itself). For this rea-
son, the functionally determinative properties of the components are provided with
tolerances. If a characteristic value lies outside of the permissible tolerance, it is
defective. Thus important functional characteristics must be tested either already
during manufacturing or at the end of manufacturing. Important disturbance factors
which must be taken into consideration as possible causes of defects are: distur-
bances caused by static forces, such as deflections effected by the workpiece weight
or clamping errors; disturbances caused by dynamic forces which lead to either
self-starting or forced oscillations; disturbances caused by thermal influences, such
as process heat or internal sources of heat in the machine tool; and disturbances
caused by tool wear (Fig. 2.1). This group also includes disturbances resulting from
the engagement kinematics between the tool and the workpiece, such as generated
cut deviations occurring during hobbing or diffraction effects on optical surfaces
caused by systematically applied tool marks.
Functionally determinative quality characteristics can be defined through macro-
geometrical properties. Macro-geometrical properties include, for example, the
accuracy of dimension, form and position of geometrical elements. Micro-
geometrical parameters include, for example, characteristic surface values, such
as roughness, mean roughness value and polishing depth. However, the functional
integrity of the process requires that certain properties are tolerated although they
lie below the surface in the peripheral surface zone. These may include hardness
values, residual stresses or certain structural properties.
The fluctuations of characteristics occurring in every manufacturing process can
basically be derived from the influence of either systematically or randomly occur-
ring disturbances. Systematic errors are caused by the system. They are reproducible
under identical boundary conditions and change the characteristic in a single direc-
tion; they cause manufacturing trends. Systematically acting disturbances cause
deviations in characteristics. If the causes of systematic errors are known, they can
3F. Klocke, Manufacturing Processes 1, RWTH edition,
DOI 10.1007/978-3-642-11979-8_2, C© Springer-Verlag Berlin Heidelberg 2011
4 2 Metrology and Workpiece Quality
Prozess
Statical disturbances
• Clamping fault
• Deflection
•
Thermal disturbances
• Internal heat sources
• Strain
• Process heat
• Environmental effect
•
Tool wear
• Displacement of cutting edge
• Flank wear
• Flank roughness
• Grinding wheel wear 
•
Dynamic disturbances
• Self-starting oscillation
• Forced oscillation
e.g. triggered over basement
•
• • •
• • •
• • •
• • •
Fig. 2.1 Variable disturbances in manufacturing
be corrected and compensated. Examples of systematic errors are geometric errors
of machine guiding elements and the development of cutting edge offset through
the continuous wear of turning tools used in turning processes. Changes of the cut-
ting force caused by sudden cutting depth fluctuations and the workpiece deflections
associated with them may also be of systematic nature (Fig. 2.2).
Randomly occurring disturbances lead to changes to characteristics, the effects
of which cannot be predicted in a strictly deterministic way. They influence the
result in both a positive and a negative direction. Thereby, the work result becomes
unstable. The influence of random disturbances on the work result can, however,
be statistically recorded. Examples of this include fluctuations in the structure of a
material, random temperature fluctuations or sudden self-starting oscillations (rat-
tling oscillations). Within real processes, it is not always easy to clearly decide
between systematic and random influences.
An important parameter describing the effect of random disturbances on a quality
characteristic (i.e. on the variation of a quality characteristic) is process capabilityor the process capability index (Fig. 2.3).
These process capability indices can only be applied if the variation of the
characteristics are normally distributed. In practice, it is frequently required that
cp ≥ 1.33. The process capability index is always taken as the dominant process
2.1 Manufacturing Disturbances and Manufacturing History 5
N
om
in
al
 d
ia
m
et
er
Pa
ss
iv
e 
fo
rc
e 
F P
 
 
/ N
 
D
im
en
si
on
al
 e
rr
or
/ µ
m
Feed direction
Cutting path
600
400
200
10
5
0
Workpiece material
Cutting tool material
Cutting speed 
Feed
Ck55 N
HW P10
Cutting part geometry:
κr = 70°
εr = 90°
rε = 0.8 mm
Cutting path
800
1000
v = 160 m/min
= 0.25 mm
vc
f
γ0 = 6°
α0 = 5°
lS = 0°
Fig. 2.2 Influence of passive force on dimensional accuracy
design criterion when large numbers of pieces must be manufactured with the high-
est possible level of productivity. Variation in manufacturing batches occur, when
it can not be guaranteed that all active elements work securely during the process
being at its performance limits. If, for example, cutting edges are randomly failed
during a cutting process because of cutting edge fractures or layer failures, this
can have an immediate, significant on the process capability. Another example of
critical process conditions is when the quantities of heat entering the workpiece
at high process performances cause thermal strain to induce an increased variation
of geometrical characteristics. These examples reveal that measuring and testing
procedures in manufacturing are necessary both to guarantee the quality of the
manufactured part and to start and develop processes.
In general, several manufacturing steps are necessary to manufacture a part.
Manufacturing progresses result from a manufacturing chain or a manufacturing
sequence. With every manufacturing step, characteristic changes (geometry, surface,
peripheral surface zone) are left behind on the part. Thus a manufacturing history
develops (Fig. 2.4). The output properties of the part following a certain manufac-
turing step in turn become the input factors for the processing steps which follow.
There can also be a significant interaction between the current step in the process
and previous changes.
6 2 Metrology and Workpiece Quality
s 3s
Process capability indices
cp = 0,71
cpk = 0,71
cp = 2,50
cpk = 1,00
cp = 1,67
cpk = 1,33
cp = 5,00
cpk = 5,00
LTL 10
15
20UTL
NV
x - LTL
x x
Mean value of the random sample UTL: Upper tolerance limit 
Estimation of the spread of the totality LTL: Lower tolerance limit
Mean deviation of the random sample NV: Nominal Value
σ :
s :
x :
Minimum proximity to process limit
Half process spread
=cpk =
Tolerance band
Process spread
=cp =
UTL - LTL
2 ( 3·σ )
xx
_
_
_
_
_
_
_
_
_
_
_
_
_
_
^
^
_
_
|x - NV|min
3 σ^
^
=
6s
T
_
Fig. 2.3 Process capability indices
Lu
Ll
Lu
Ll
Machining
center
Grinding
machine tool
k1a c2b knx
Turning
milling Hardening
Hard
machining Assembling
Forging,
cold
extrusion 
Enveloping cut
deviation Texture
Lu
Ll
Anisotropism
Lu
Ll
Lu,l Upper and
lower limit
c1a k2b cnX
k, c Damping and
stiffness
Machining
center
Geometry
Fig. 2.4 Manufacturing history (Source: Getrag Ford)
2.1 Manufacturing Disturbances and Manufacturing History 7
Box part: Automobile-gearbox made of Mg-alloy 
AZ 91hp
Standard tolerance IT 4
Requested cpk = 1.67
Manufacturing process: Primary shaping, machining
Manufacturing dispersion: Affected by residual stress, chucking 
0.008
0.01
Fig. 2.5 Representative case study – box part
The hole in a transmission part shown in Fig. 2.5 could not be manufactured in
a procedurally secure way with the required process capability index and specified.
There can also be a significant interaction between the current step in the process
and previous changes. The hole in a transmission part shown in Fig. 2.5 could not
Work piece:
Blisk
(Blade integrated Disk)
made of TiAl6V4
Tool:
End milling cutter
(D = 8 mm)
Manufacturing process: Forming, 5-axis-milling
Manufacturing dispersion: Process instability, chatter marks, long 
and slander tools 
Source: MTU
stabil
instabil
stability boundD
ep
th
 o
f c
ut
/m
m
Speed / min–1
Fig. 2.6 Stability bounds when milling a BLISK
8 2 Metrology and Workpiece Quality
Manufacturing process: Forming, turning, drilling, broaching
Manufacturing dispersion: Process anomaly,
tool life decrease of the part
Source: Rolls-Royce
Work piece: Turbine disk
made of 
Inconel718
Tool: VHM-drill 
(D = 8 mm)1000 10000 100000 1000000
Numbers of change of load
K
t xσ
ap
pl
. / 
M
Pa
50 µm
Without anomaly 
Smeared material
Low thermal influencing
High thermal influencing
Fig. 2.7 Surface damages when drilling with a worn drill
be manufactured in a procedurally secure way with the required process capability
index and the demanded productivity. This is due to released residual stresses which
arised in the previous die-casting process.
Figure 2.6 shows a part manufactured by milling with long, slender end mills.
The milling process is only stable for certain combinations for cutting speed
and depth of cut. When these areas are exceeded, rattling oscillations occur. These
characteristic stability areas must be known; otherwise the process cannot be carried
out at the performance limit.
However, not only impermissible macro-geometrically or micro-geometrically
recognizable changes may limit the process, but also impermissible changes in the
peripheral surface zone. Figure 2.7 shows a turbine disk which was exposed to
high mechanical and thermal stresses caused by machining the drill holes. These
stresses damaged the peripheral surface zone in a way that significantly diminished
the fatigue strength of the part.
2.2 Measuring and Testing
Quality characteristics must be tested during the manufacturing due to the boundary
manufacturing conditions mentioned above. This can be done by means either of
samples or of a 100% test. A 100% characteristic check is prescribed for many
safety-related parts. Though, a 100% test does not assure freedom of errors.
2.2 Measuring and Testing 9
Testing means determining whether the test sample exhibits the required char-
acteristics, such as dimension, form and surface quality. This determination can be
made in different ways, thus the distinction between subjective testing and objective
testing.
Subjective testing is performed by means of the sensory perception of the tester,
e.g. visual and tactile tests. Examples of this are the evaluation of burr formations
on the edges of parts and of shadings on polished surfaces.
In objective testing, the tester is supported by testing equipment. There are two
further methods which fall under this category; measuring and gauging.
Measurement is the comparison of a characteristic with a measurement standard.
The result is a measured value. Gauging is the comparison of the test object with
a dimension or a form. The result is information regarding whether the specified
requirements were satisfied or not (go or not-go).
Testing equipment is subdivided into measuring instruments, gauges and aids.
Measuring instruments and gauges are distinguished by material measures cor-
responding to the measurand. Material measures can be of mechanical, electric,
optical and electronic nature. Mechanical material measures may be, for example,
distances between surfaces or angular positions of surfaces. Indicating measuring
instruments have movable markers, divisions or counters. When using these devices,
the measuredvalue reading can be either direct analogue or digital. In contrast to
this, the gauge body corresponds merely to the dimension or the dimension and form
of the required characteristic. Intermediate values cannot be defined.
Table 2.1 provides a survey of the attainable accuracy values when using different
manufacturing processes.
Table 2.1 Survey of manufacturing qualities
16
Primary shaping
Die forming (warm)
Extrusion (cold)
Turning 
Drilling
Reaming
Milling
Grinding
1 2 3 4 5 6 7 98 10 11 12 13 14 15
Standard tolerance IT
Deep drawing
Erosion
2.2.1 Measurement Errors
As a matter of principle, every measurement result is subject to error. Possible
reasons for measurement errors are shown in the ISHIKAWA diagram in Fig. 2.8.
The causes of measurement errors may lie in the measuring method, the measuring
10 2 Metrology and Workpiece Quality
Measurement
Measuring method Preparation Environment
Temperature
Humidity 
Air pressure
Vibration 
Extraneous light
Interference field
Measuring personMeasuring deviceMeasured object
Preliminary
arrangements
Cleanliness
Protocol
Calibration
Unconsidered
disturbances
Aptitude 
Abbé principle error
Testing
deviation
Aim of the
measurement 
Measuring uncertainty
Aptitude
Measuring forces 
Assembling
Aptitude 
Routine 
Evaluation
Visual acuity
Mistake
Attention
error
Fig. 2.8 Causes for measuring errors by the measuring procedure
instrument, the measured object, the person measuring, the preparation involved and
the environment.
In order to achieve a high measurement accuracy, it is necessary to take the
abovementioned influences into consideration and, where applicable, to minimize
them, compensate them, or at least to estimate them in terms of their magnitude.
It can principally be distinguished between systematic and random disturbances in
measurement procedures, as well [DIN1319a].
2.2.2 Macro- and Microgeometry of Components
2.2.2.1 Structural Deviations
With regards to components, the distinction is often made between macro-
geometrical parameters and the surface quality. Macro-geometrical parameters refer
to deviations of dimension, form and position. The surface quality is defined by
roughness parameters. The transitions between these categories are not always
clearly definable. DIN 4760 offers a general system for organizing structural devi-
ations (Fig. 2.9). Therefore it is necessary to define the term “surfaces” at the
beginning. The real surface (primary surface; see DIN EN ISO 4287) is the sur-
face actually present on the part. The actual surface is the metrologically registered
surface. It may differ from the real surface, since every measuring method can only
approximate the real surface. A geometrically ideal surface is assumed in designs
2.2 Measuring and Testing 11
Superposition of structural deviations
1st up to 4th Order
5th Order: textural structure
6th Order: lattice structures of the material
4th Order: scores, scales (roughness)
3rd Order: grooves (roughness)
2nd Order: waviness
1st Order: form deviations 
Structural deviations
(in superelevated representation)
Not easily representable
Fig. 2.9 Structural deviations, acc. to DIN 4760
and forms the basis of tolerances. In Fig. 2.9 six orders of structural deviations are
defined on the basis of these observations.
Structural deviations of the 1st order (see also the following sections) are fre-
quently the result of systematic errors. With regards to waviness, i.e. the structural
deviations of the 2nd order, one cannot clearly define whether they are caused by
systematic or random influences. The unbalance of a rotating tool and any peri-
odical oscillations caused by it are forced, while sudden rattling oscillations are
self-starting. In general, fundamentally different actions must be implemented in
order to exclude any systematic or random causes of error. Structural deviations
of the 3rd order also occur regularly. They are to be attributed to the penetration
between tool and workpiece and are often determined by means of penetration
calculations. Examples of these are kinematic roughness associated with turning,
surface marks created in peripheral milling and generated cut deviations created
in hobbing. In such cases, the structural deviations can be influenced in a targeted
way by means of generation kinematics and tool design. The higher orders of struc-
tural deviation are primarily random in their occurrence. Examples of structural
deviations of the 4th order include chip formation processes and removal processes.
Roughness of the 5th order is rendered visible by structural properties on the surface.
This can play a significant role in the high-precision machining of metallic optical
mirrors. Thus in high-precision turning of multicrystalline metals, grain boundaries
may become visible because the individual crystals exhibit varying orientations and
therefore varying stiffnesses. In this case, anisotropism of the grains becomes visible
on the surface.
12 2 Metrology and Workpiece Quality
In general, all the structural deviations on a real surface are superposed. Filters
are employed to separate roughness and waviness in a measurement process [DIN
EN ISO 4287]. The following sections will treat macrogeometrical deviations of
structure, form and position, as well as the corresponding measurement technology.
Section 2.4 and the following sections describe the metrological recording of higher-
order structural deviations.
2.2.2.2 Form Deviations
Form deviations refer to deviations from a specified ideal geometrical property,
such as straightness, evenness, roundness or cylindric form [DIN EN ISO1101,
VDI2601].
The following will introduce some examples of form errors and their causes
(Fig. 2.10).
• A cause for deviations from the cylindric form of a workpiece can be the incorrect
alignment of the workpiece on the machine tool with respect to the tool.
• Another cause for form deviations is the continuous alteration over time of the
tool geometry caused by wear.
• Thermal displacements of the workpieces, tools and machine tools can also lead
to form deviations.
• Deviations from cylinder form can also arise when the workpiece deflects
because of a radial strain. This can happen, for example, e.g. when turning long,
slender parts when the workpiece is not supported.
• Roundness errors can develop through the incorrect clamping of the workpieces
on the machine tool.
2.2.2.3 Position Deviations
Position deviations are deviations of the position of a geometrical element with
respect to a reference, such as an edge, circulation line or axis of the predetermined
position. In general, the position of two surfaces or axes in relation to each other is
indicated by simple length or angle specifications.
Fig. 2.10 Deviations from
circularity and cylinder form
2.3 Length Testing Devices 13
Fig. 2.11 Deviations from
concentricity, straightness
and parallelism
Position deviations as shown in Fig. 2.11 may arise, for example, through
incorrect clamping or defective clamping devices.
2.3 Length Testing Devices
2.3.1 Material Measures
A material measure in length measurement technology represents lengths or angles
by means of fixed distances or angles between surfaces or lines. Setting standards
are material measures. The material measures used most frequently are gauges.
The most important and most precise material measures in length measurement
technology are parallel gauge blocks. They consist of two plane-parallel measuring
surfaces made of hardened steel or cemented carbide. Parallel gauge blocks are used
in prismatic and cylindric forms. Gauge blocks are composed of grades. A standard
set contains 5 grades each containing 9 blocks, from which almost any dimensions
can be assembled (see Fig.2.12). The measuring surfaces are superfinished and even
enough to allow, after careful cleaning of the measuring surfaces, individual gauge
blocks to be joined together so that they adhere.
Fig. 2.12 Gauge blocks (Source: Mahr)
2.3.2 Gauges
A gauge measures dimensions or forms generally according to limit dimensions
[DIN2257]. It can be distinguished between inspection gauges, limit gauges and
form gauges (Fig. 2.13).
14 2 Metrology and Workpiece Quality
21,5
G
o
G
u
Fig. 2.13 Types of gauges: inspection gauges, form gauges, limit gauges
Square profile
Go side
Not go side
Multitooth profile 
Go side
Not go side
Go side
Not go side
Shaft
Go
side
Not go side
Drill hole
Fig. 2.14 Different types of limit gauges
Inspection gauges are parts of a gauge set for which the measure is assembled
from the combination of gauges. Examples are the combination of parallel gauge
blocks and the setting of feeler gauges.
Form gauges allow for the testing of profiles using the light-slit method. They
include angles, radius gauges and thread gauges.
Limit gauges function according to Taylor’s principle: The go gauge must be
designed in such a way that the dimension and form of the workpiece can be
inspected when combined with the gauge. Not-go gauges are only used to inspect
single dimensions.
Limit gauges used most frequently are calliper gauges, limit plug gauges, gaug-
ing rings and thread gauges (Fig. 2.14). Gauges are the simplest testing equipment
used in industrial manufacturing.
2.3.3 Indicating Measuring Instruments
Indicating measuring instruments consist of a material measure and a display mech-
anism. The indicating range is the range of measured values that can be displayed
on a measuring instrument. The measuring range is the part of the indicating range
for which display deviations lie within specified or agreed limits. The suppression
range of a measuring instrument is the range which must be exceeded in order that
2.3 Length Testing Devices 15
the instrument can begin to display values. The hysteresis of a measuring instrument
equals the difference of the display values determined by measuring from different
directions. The hysteresis of a measuring instrument is often not constant, which
means that only a certain lower limit is indicated. The reversal error is principally
a systematic error. It can be compensated. The sensitivity of a measuring instru-
ment is the ratio of an observable change on the measuring instrument display to the
causative parameter. Using length measuring instruments, the sensitivity is defined
as the ratio of the range of the indicating element (e.g. the indicator) and the range
of the measuring element (e.g. the spindle or measuring arm). Conversion between
the measurement range and the display may be based on different physical prin-
ciples. The most frequently used conversion elements are mechanical, electrical,
pneumatic, optical and electronic conversion elements.
2.3.3.1 Instruments with Mechanical Converters
Instruments with mechanical converters utilize mechanical or optical indicating
elements. In length measurement technology, the most often used measuring instru-
ments function on the basis of mechanical converters, e.g. threads, racks, gearwheels
and blade segments (Fig. 2.15).
With mechanical callipers with sliding glideways as the conversion element, the
material measure is mounted on the bar on which the display, i.e. the slider, is
moved. Mechanical callipers have two different scales which can be read off against
each other by the observer. The scale on the slider is called a vernier (Fig. 2.15).
In the case of micrometers, highly precise threads are used as converters and as
a material measure. The system comprises a measuring spindle running in a sleeve,
whose forward face is designed as the measuring surface and whose back end car-
ries a ground thread as material measure. Typical thread pitches are 0.5 and 1 mm.
As seen with callipers, two different scales are read off against each other. One scale
is mounted on the sleeve, the other on the scale drum. In order to avoid influencing
Fixed measuring jaw
Moveable
measuring jaw
Clamp
Bar
Vernier
Linear scale
Measuring
ambos
Measuring
spindle
Scale sleeve
Thimble
Coupler
Fig. 2.15 Setup of callipers and micrometers
16 2 Metrology and Workpiece Quality
Arm with
blade
segment
Rack
Pinion
Prestressed
spring
Plug gauge
Fig. 2.16 Setup of dial gauges and dial comparator
the measurements through large fluctuations in the measuring force, the latter is lim-
ited to 5–10 N by means of a coupler (ratchet coupling). Micrometers are designed
both as outside micrometers for external dimensions and as inside micrometers for
internal dimensions (Fig. 2.15).
Dial gauges are length measuring instruments equipped with racks and gear-
wheels as conversion elements which allow a larger view of the path of the plug
gauge. The material measure lies in the gear mechanism, the conversion element
which leads to measurement reversal errors due to anisotropic friction and potential
tolerance. Needle positions greater than 360◦ are possible (Fig. 2.16).
Dial comparators are the most precise mechanical length measuring instruments.
They are equipped, unlike dial gauges, with a gear mechanism as lever arm system,
gear segment and pinion as conversion elements. Thereby, the movement of the plug
gauge is transmitted to the needle. The design only allows needle positions smaller
than 360◦ (Fig. 2.16).
In principle, all the instruments with mechanical conversion elements described
here can be equipped with frictionless optical indicator elements in order to improve
precision. Given constant conversion behaviour, the more precise indications the
measuring instrument should display, the longer the indicator must be. However,
since there is only a limited amount of space in the instruments and, above all,
since the inertia of the indicator can lead to distorted measurement results when the
measurement range is small, optical indicator elements are used. The principle is
basically this: a beam of light is directed to the material measure and, according to
the position of the material measure, this beam is reflected in a different direction.
The optical indicator then displays the measurement range of the material measure
without friction and inertia.
Further potential for increasing precision can be exploited by means of a fine
graduation of the scale which can no longer be observed with the naked eye, thus
necessitating optical magnifying aids in the form of microscopes. By means of the
microscope, a mechanism consisting of an objective and an ocular, a magnifica-
tion of a real image by the objective can be viewed with the ocular. In this way, a
remagnified virtual image of the object is created. The total magnification through
2.3 Length Testing Devices 17
a microscope is the product of the objective and ocular magnifications. The scale
graduation can be applied in the ocular directly or by means of an ocular plate.
Microscopes including measuring oculars are called measuring microscopes.
2.3.3.2 Instruments with Electrical Converters
These instruments are used to determine changes in the measurand caused by alter-
ations of electrical properties, such as resistance, inductance and capacity. The turns
of a coil, the distance between plates of a condensator and the resistance of an
electrical conductor can all serve as material measure. Inductance, for example, is
altered by changing the immersion depth of an iron core; the capacity of a plate
capacitor is a function of the plate distance (Fig. 2.17).
C
˙
s 
UB
UA
=
UA = Bridge-output voltage
UB = Bridge-supply voltage
C = Capacity
s = Measurement range
o
= Electric constant
r = Dielectric constantA = Surface area 
d = Distance
Position encoder
Sensor
Position encoder
Output to
measuring
amplifier 
L1
L2
s
R4
UA
UB
R3
d
C = r˙ 
Ao
˙
ε
ε
ε ε
Fig. 2.17 Capacitive and inductive measuring principles
18 2 Metrology and Workpiece Quality
Inductive position encoder
Fig. 2.18 Inductive working sensor in real execution
An alteration of the ohmic resistance can be effected by changing the conductor
length and/or by a cross-sectional variation [Hoff04, Grot05]. Figure 2.18 shows a
real execution for multi-point measuring with inductive working sensors.
2.3.3.3 Instruments with Optical Converters
These length measuring instruments exploit the wave property of light. The wave-
length of light is used as the material measure. Figure 2.19 shows the functional
principle of a laser interferometer.
Mirror
Laser
Detector
Wedge prism of
the measuring
beam
Measuring path
Polarizing
beam splitter 
Displacement
Wedge prism
of the reference beam 
R
ef
er
en
ce
 p
at
h
Interference of
measuring und
reference beam
Fig. 2.19 Longimetry with laser interferometer
2.3 Length Testing Devices 19
ϕ2
ϕ3
λ3
λ1
λ2
Time
Am
pl
.
l
Am
pl
.
Time
Am
pl
.
Time
TimeA
m
pl
.
Interference of two coherent
light waves out of phase
(destructive interference) 
Monochromatic light
Am
pl
.
Time
Interference of two coherent
light waves with the same
 phase length
(Addition of illuminance) 
λ: Wavelength
ϕ : Phase difference
Light mixture of different
wavelengths 
Monochromatic and 
coherent light
λ ≠ const. ϕ ≠ 0 λ = const. λ = const.ϕ ≠ 0 ϕ = 0
ϕ2
ϕ3 λ
λ
Fig. 2.20 Optical measuring principle, interference
A light source emits coherent, monochromatic light whose beam is split into a
reference beam and a measuring beam at a polarizing beam splitter. These beams are
reflected on a respective wedge prism and combined again through superposition.
Finally, the combined beam is detected and evaluated. If the wedge prism of the
measuring beam is shifted, interferences occur because of the difference of the opti-
cal paths of the two partial beams after combining them. These are fluctuations in
light intensity via eliminating and amplifying light. Integral multiples of the wave-
lengths lead to an amplification, phase differences of 180 degrees to an elimination
of light (Fig. 2.20).
The number of interference lines is directly related to the path being measured.
Since the wavelength of the light is a function of temperature and also air pressure
and humidity, these influences are compensated within the device. The advantages
of such systems lie in their large measuring range and especially in their contactless
measurements.
2.3.3.4 Instruments with Pneumatic Converters
These instruments use pressure and flow rates as material measures (Figs. 2.21
and 2.22).
This measuring principle can be used for both tactile and contactless measure-
ment (Fig. 2.22).
There are two types of such measuring instruments: high-pressure and low-
pressure instruments. High-pressure instruments work with an operating pressure
greater than 0.5 bar, low-pressure instruments with an operating pressure lower than
20 2 Metrology and Workpiece Quality
1 2 3
4
s
Volume measuring
method
1. Air filter
2. Pressure control
3. Flowmeter
4. Measuring nozzle
1 2
h
6
3 4
5
s
Velocity measuring
method
1. Air filter
2. Pressure control
3. Venturi-nozzle
4. Flow of valve
5. Measuring nozzle
6. Differential pressure
 manometer
Fig. 2.21 Measuring principle of pneumatic distance measurement
Air inlet
Valve
Follower
Work piece
Guide body
Measuring
nozzles
Drill hole
Tactile measurement Contactless measurement
Fig. 2.22 Pneumatic sensor
0.1 bar. The pressure range between 0.1 and 0.5 bar is outside the operating range
of the instrument [DIN2271a].
Figure 2.23 shows the functional principle of contactless pneumatic measure-
ment. The left side shows an external measurement, the right side an internal
measurement. If the measuring rod is designed in a way that both measuring nozzles
2.4 Surface Inspection 21
Measuring
nozzle ∅dm
h
Annulus
pa pa
ps
Work piece surface Measuring
nozzle 
Test sample
Fig. 2.23 Functional principles and examples for pneumatic measurement
work within the linear range of the characteristic diagram, then an exact centring in
the hole is unnecessary, since the air currents of the two nozzles are added and the
sum represent the measure for the total gap width. Since the type of hole measure-
ment described is a two-point measurement, deviations from hole roundness and
diameter can be determined by turning the rod or the workpiece.
The measuring rods are often specially adjusted to the measurement of a spe-
cific workpiece. Therefore, for economical reasons, pneumatic measurement is used
predominately in serial production, especially if a 100% test is required.
The advantages of contactless pneumatic measurement lie in the self-cleansing
effect of the measuring device (i.e. the workpieces do not generally have to be
cleansed of oil, dirt or micro-chips) and in the quickness of the measuring process.
2.3.3.5 Electronic Measuring Instruments
These devices work with photoelectric sensors, by which light-dark fields of a gauge
are converted into electric signals. The material measure is realised through the
light-dark fields. Impulse gauges and code gauges are preferably used as path mea-
surement systems. To be precise, it is a mechanical guidance system whose positions
are detected and indicated optoelectronically. This path measurement system is also
executed with callipers, dial gauges and dial comparators. It is also realized in
coordinate measuring devices and machine tools. To control measuring axes and
to evaluate data, an exact control system and effective software are required which
can influence the quality of the measuring result and the applicability of the data
[Pfei01]. The measurement of freeform surfaces and tooth-flank topographies on
spur gears and bevel gears with coordinate measuring devices is currently state of
the art (Fig. 2.24).
2.4 Surface Inspection
The task of characterizing a technical surface consists in attaining a complete
topological record of its three-dimensional geometry and describing it. Often, in
order to simplify the measuring process, only parameters for surface roughness
along a single measured length are registered (one-dimensional parameters). This
22 2 Metrology and Workpiece Quality
Fig. 2.24 3-D coordinate measuring machines
procedure is frequently followed in practice and is sufficient in many cases. For cer-
tain applications, e.g. in optics or for characterizing grinding wheel surfaces, more
comprehensive, multi-dimensional surface descriptions must be used. However, the
following will exclusively treat one-dimensional parameters.
2.4.1 Surface Parameters
In most of the procedures used in industrial surface inspection technology, only
structural deviations of the second or higher orders are analyzed and measured on
surface sections. These sections must be statistically representative for the entire
surface [DIN2257]. The surface can be detected by means of surface sections or
the bearing surface (Fig. 2.25) [DIN4760]. Profile sections are sections which are
oriented to the surface normally, tangentially or at other angles.
First, a profile section will be used to illustrate some of the basic terms of surface
inspection technology (Fig. 2.26). With respect to a profile, the distinction is made
between the test length Lt of the surface section being metrologically detected and
Tangential cross sectionProfile
α
Profile
Cross section parallel to
the tangentialcross section Diagonal crosssection
Vertical cross sections
Fig. 2.25 Registration of structural deviations by surface cuts
2.4 Surface Inspection 23
Ra
Z (x)
Measured length L
Zt
1
L1 L2 L3 L4 L5 L6
Zp
1
Zv
1
Zv
2
Zp
6
x
Fig. 2.26 Fundamental terms of surface inspection technology
the measured length L, which is used for evaluation (L < Lt) [Grot05]. The detected
profile (actual profile) of a surface depends on the measuring procedure and the filter
used, thus it only represents an approximate image of the actual surface. The refer-
ence profile shifted within the measured length perpendicularly to the geometrically
ideal profile is defined as the middle profile. This is oriented in such a way that the
surface areas above and below the middle profile line are equally large (Fig. 2.26).
By using different measuring procedures and depending on which filter is used,
different profiles can be determined [DIN EN ISO 4287]. These are the P-Profile
(primary profile), the R-Profile (roughness profile) and the W-Profile (waviness
profile) (Fig. 2.27).
Grenzwellenlänge: λ = 0,75 mm
1 
µm
250 µm
1 
µm
250 µm
1 
µm
250 µm
W-Profil (Welligkeit)
R-Profil (Rauheit)
Without wave filter
High-pass filter
Low-pass filter
W-Profile (waviness)
R-Profile (roughness)
Cutoff wavelength: λ = 0,75 mm
250 µm
1 
µm
250 µm
1 
µm
250 µm
1 
µm
P-Profile (waviness and roughness)
Manufacturing technology: Short-stroke honing
Average roughness: Ra = 0.75
Scanning system: HAT 25/6
Scanning path: 10 mm
Fig. 2.27 Separation of waviness and roughness by wave filter
24 2 Metrology and Workpiece Quality
The examples given below refer to the R-Profile (roughness) (Fig. 2.26).
According to DIN EN ISO 4287, the following roughness parameters can be
distinguished:
• The height of the highest profile point Rp: value of the y-coordinate Z(x) of the
highest profile point of the middle profile line within the sampling length Li.
• The depth of the deepest profile valley Rv: value of the y-coordinate Z(x) of
the deepest point of the profile of the middle profile line within the sampling
length Li.
• The total height of the profile Rt: the sum of the highest profile point and the
depth of the deepest profile valley within the measured length L.
• The greatest height of the profile Rz: the sum of the height of the highest pro-
file point Rp and the depth of the deepest profile valley Rv within a sampling
length Li.
• The mean roughness value Ra: the arithmetic mean of the values of the
y-coordinates Z(x) within a sampling length Li.
Ra = 1L
L∫
0
|Z(x)| dx (2.1)
Horizontal parameters (distance parameters) are also designated as bearing
lengths. They are determined by means of tangential sections. Forming a ratio
of of the summed single bearing lengths and dividing by the measured length
yields the relative bearing length (material ratio) in a specified section depth c.
By creating sections at different depths c, the bearing ratio curve – also called the
ABBOTT-FIRESTONE curve – can be determined (Fig. 2.28).
The bearing ratio of the roughness profile Rmr(c) is calculated as follows:
Rmr(c) =
∑
Xi
L
(2.2)
Measuring length L
c
10060200
Rmr (c ) / %
x1 x2 x3 x4
x5 x6 x7 x8 x9 x10 x
Z(x)
Fig. 2.28 Bearing ratio curve of the profile, acc. to DIN EN ISO 4287
2.4 Surface Inspection 25
The bearing ratio can be calculated on the basis of the primary profile (Pm), the
roughness profile (Rm) and the waviness profile (Wm).
Using the average roughness, Table 2.2 provides an overview of the surface
roughness values which can be achieved with different manufacturing processes.
With respect to general specifications on achievable surface values, one must
consider the fact that it is not necessarily possible to deduce the manufacturing pro-
cess when using one-dimensional parameters for describing the surface [Abou76].
Due to the characteristic engagement conditions between the workpiece and the tool
one would also have to specify, for each surface parameter, which manufacturing
process is to be used to create that parameter. This problem is somewhat alleviated
if multiple one-dimensional surface parameters are used instead of just one when
describing the manufactured surface or if surface reference standards are available
(Figs. 2.29 and 2.30).
Table 2.2 Achievable average roughness
10
00
Primary shaping
0.
1
0.
25
0.
4
1 2
.5 4 1610 25 40
0
25
0
Achievable average roughness Rz / µm 
0.
04
16
0
30
0
63
0
Die forming
Extrusion
Turning
Drilling
Reaming
MillingMilling
Grinding
Erosion
Rough and finish turned
Planed
Honed
Grinded
Fig. 2.29 Surface quality for several manufacturing operations
26 2 Metrology and Workpiece Quality
Flat
lapping
Reaming
Flat
grinding
Horizontal
milling
Face
milling 
Cylindrical
turning
R
ou
gh
ne
ss
 R
t (µ
m)
Ar
ith
m
. a
ve
ra
ge
ro
u
gh
ne
ss
 R
a 
(µm
)
10
3.
0
0.
55
0.
05
0.
1
0.
2
0.
4
0.
8
1.
6
2.
5
8.
0
4.
0
16
32
50N
7
N
6
N
5
N
4
N
3
N
2
N
10
N
9
N
8
N
7
N
6
N
5
R
ou
gh
ne
ss
 m
ar
kl
IS
O
 R
 1
30
2–
19
71
Su
rfa
ce
 m
ar
k
D
IN
 1
40
Ar
ith
m
. a
ve
ra
ge
ro
u
gh
ne
ss
 R
a 
(µm
) 1
2.
5
6.
3
3.
2
0.
8
0.
4
R
ou
gh
ne
ss
 R
t (µ
m)
6,
0
1.
0
1.
6
1.
6
Fig. 2.30 Surface reference standard (Source: Rupert)
2.4.2 Subjective Surface Inspection
In some cases, a visual inspection can determine whether the processed surface
meets the specified requirements. A testing device frequently used in the work-
shop is the straight edge, which is used to check the evenness of a surface using
the light-slit method. In workshop operations, spotting straight edges or spotting
plates are used to detect form deviations or contact patterns on the workpiece surface
(Fig. 2.31).
Impression methods using lacquers have been developed for surfaces which are
difficult to access (e.g. drill holes, internal surfaces). The lacquer is applied onto the
Spotting straight edge
Spotting medium
Work piece
Grey gleaming
areas
(bearing points) 
Light areas 
(lowest points) 
Black areas 
(spotting medium) 
Spotting plate
Fig. 2.31 Surface inspection by spotting
2.4 Surface Inspection 27
test surface and pulled off after drying. The contact patterns attained on the lacquer
film thus attained can then be evaluated.
A frequent use of spotting is detecting contact patterns in quality testing and in
the assembly of gear teeth (especially bevel gear teeth). Applying spotting paste
and then rolling the teeth under light strain causes the bearing points on the tooth
flanks to become visible. These bearing points can then be compared to the reference
contact pattern. Nowadays in individual cases machine-tool guiding elements also
become scraped, therefore, spotting is used to judge the surface.
2.4.3 Surface Measurement
All the physical principles mentioned in the context of length measurement technol-
ogy are also fundamentally applicable to surface measurements. Since mechanical
and optical measuring methods have proven to be the most effective in this field, the
following discussion will limit itself to these methods.
2.4.3.1 Mechanical Measuring Methods
Devices for surface measurement which function on the basis of a mechanical
functional principle are generally referred to as stylus instruments. Mechanical mea-
suring methods are classified as either scanning methods or sensing methods. In the
scanning method, a contact stylus descends upon the surfaceto be tested at a speci-
fied frequency. The surface is guided under the needle with a constant feed rate. The
path of the contact stylus can be visualized mechanically, optically, electrically or
electronically. The contact stylus can either be raised to a fixed level (the WOXEN
principle) or raised by a fixed amount from the respective point of impact to the
surface (differential tactile procedure) (Fig. 2.32).
In the differential tactile procedure, the impact energy of the contact stylus is
lower in comparison to the WOXEN principle and exhibits marginal dispersion. As a
result, the penetration depth of the needle remains constant, the measuring accuracy
being thus higher than in the case of the WOXEN principle. In the case of instruments
functioning according to the sensing method, the contact stylus is guided continu-
ously over the surface. The needle rises and falls in line with the profile pattern
Scanning method Sensing method
Woxen principle Differential tactile procedure Differential tactile procedure
Fig. 2.32 Mechanical measuring methods
28 2 Metrology and Workpiece Quality
Caliper
Caliper slide
Table
Plane of reference
system 
Caliper slide
Caliper
Skid
Oscillating scanning
system 
Caliper
Scanning
arm 
Skid
Semi-rigid
System 
Fig. 2.33 Different surface scanning systems
(Fig. 2.32). The lifting motion is indicated relative to a reference point defined in the
device or to a reference level. Here too, mechanical, electric, optical and electronic
converters are used.
Sensing profile methods are the most widespread in practice. Independently of
the design type of the contact stylus instruments used, three system designs are
distinguished (Fig. 2.33).
Plane of Reference System
In this system, the scanning unit is guided on a reference surface (plane, cylinder)
which corresponds to the ideally geometrical surface of the test sample and is ori-
ented along the surface to be measured (Fig. 2.33). Aside from errors arising due to
the calliper geometry, this scanning system provides a faithful transmission of the
roughness and waviness values of the test sample. When measuring small or very
large surfaces, however, the handling of this scanning system can frequently become
unwieldy. An alternative is the reference surface contact system. Here, the work-
piece is conveyed on very precisely guided slides in a horizontal direction beneath
the firmly anchored scanning system. Besides roughness, the macrostructure of a
surface can also be detected in certain areas.
Semi-Rigid System
In semi-rigid systems, the scanning unit, which contains the contact stylus, is guided
on the skid gliding on the surface to be measured (Fig. 2.33). This system has the
advantage that it requires little space and is thus suitable for measuring small or
hard-to-access surfaces. A disadvantage is that the system requires an orientation to
the surface to be measured which can cause parts of the profile to be transmitted in
a distorted manner.
Oscillating Scanning System
Two glide skids guide the scanning unit of the oscillating scanning system
(Fig. 2.33). The latter is oriented towards the surface to be measured and is thus
2.4 Surface Inspection 29
comfortable to operate. However, it requires more space than the single skid system
and can therefore not be used for small or hard-to-access surfaces. The distortions
caused by long-wave profile sections are smaller than those in half-rigid systems,
because the skids are flat and farther apart. Nevertheless the waviness must be
filtered out in many cases [Henz68].
Usually, the magnification gauges for x- and y-coordinates are selected in a highly
varying way in profile records. This is necessary because the measuring lengths
(abscissa) lie in the mm region and roughness parameters are represented on the
ordinate which lie in the μm region. The optical impression of the profile record is
thus strongly distorted in comparison to reality. The contact styluses used in con-
tact stylus instruments often have an apex radius of 2 μm. They function with a
bearing strength of 0.5 N, which can result in considerable surface pressures (up
to 6000 N/mm2), which possibly cause alteration of the test surface. On the other
hand, excessively large contact stylus radii distort the result; they act like mechan-
ical filters. Thus the optimal conditions must be determined on the basis of the
material of the test sample and documented in their entirety in the measurement
report.
2.4.3.2 Optical Measuring Methods
White-Light Interferometer
This technology differs from length measuring technology both in that it employs
white light, i.e. light with the entire wavelength spectrum, and in that it does not just
use one beam, as with laser interferometers, but rather an entire bundle. A reflected-
light microscope is used to display an image of a section of the test object on a
detector (e.g. CCD camera). By using different interference lenses, a beam splitter
can be used to superimpose a highly accurate reference surface with the image of
the test object on the same scale. The interferences which are then created can be
detected and evaluated (Fig. 2.34).
The topography of the test object creates a spatial modulation of the light
intensity in the interference image. Depending on the surface to be measured,
two different measuring modes are used: one to characterize very flat surfaces
with an average roughness value Ra < 1 nm and one for all other surfaces. Steps
and roughnesses of up to several millimetres high can be displayed using this
technology.
Fringe Projection
The fringe projection method functions according to the triangulation procedure, in
which equidistant stripe patterns are observed and evaluated at a certain angle, i.e.
the triangulation angle. The stripe patterns on the test object are detected and evalu-
ated from a certain position, with the projected stripes following the arbitrary form
of the test surface. These stripes appear from the observer’s standpoint to be “inter-
ferences”, though they only appear this way because the location of the projection
30 2 Metrology and Workpiece Quality
Lenses
Light
Detector
Displacement
Interference of
measuring and reference
beam 
Te
st
 o
bje
ct
z
x
Reference mirror
D
is
pl
ac
em
en
t
Fig. 2.34 Principle of a white-light interferometer (two-beam interferometer)
surface points varies in relation to that of the projector. The stripe patterns then are
evaluated interferometrically on the basis of fluctuations in light intensity. The most
important geometrical values on which this method is based are:
• the real distance of the projected stripes
• the stripe distance registered from the position of observation
• the triangulation angle
Fringe projection is used, for example, to judge surfaces bent over a large area,
such as those found in deep-drawing tools and on deep-drawn parts. The advantages
of this method are a high measurement speed and a spatial or laminar scanning.
A disadvantage is that the test objects may not be transparent or reflective.
2.5 Inspection of the Workpiece Rim
The functional behaviour and applicability of a component depend not only on its
macrogeometry and surface roughness, but also from the physical properties of the
material both on the interior and near the surface. While the inspection of material
properties properly belongs to the field of materials science and materials testing
and thus cannot be treated here in further detail, the following will discuss some
of the properties of technical surfaces and of layers near the surface (i.e. the rim
zone) and methods used to measure rim zone properties. Technical surfaces can be
categorized according to types of load into the following groups:
2.5 Inspection of the Workpiece Rim 31
• external surfacesof technical products of all types, i.e. visible surfaces, covering
surfaces, indicating surfaces, etc. They are generally mechanically unstrained,
but are exposed to climatic or environmental stresses.
• surfaces subjected to heat, radiation or electrical currents, such as insulating sur-
faces, electrical contacts, or the like. Such surfaces are referred to as thermally
stressed, radiation stressed or electrically stressed surfaces.
• surfaces which come into contact with fluids or gases. On the one hand, a
corrosive stress may be predominant, or a flow stress with cavitation and ero-
sion processes may occur in the case of flowing media. Also, surface boundary
currents may be influenced via microstructures (ribblets).
• surfaces in mechanical contact with moved counter bodies. This strain is referred
to as tribological stress. These stresses are found, for example, in typical machine
elements, such as bearings, couplings, brakes, gear-teeth, etc. With this type of
stress, different kinds of wear may occur.
• optical surfaces used to form and conduct electromagnetic waves. Optical
surfaces are produced with mirrors and transparent components. The basic
beam-conducting and forming principles are reflection, refraction and diffraction.
• biologically stressed surfaces exposed to the effect of microorganisms.
2.5.1 Surface Layers
The properties of technical surfaces relevant to component behaviour are determined
through the entirety of the physical and chemical properties of the surface layer.
These properties include textural structure, hardness, strength and residual stresses
in the rim zone near the surface. The surface rim zones of technical bodies are
created through machining processes. A part of the energy used to create the surface
always flows into the workpiece and is either stored or causes change processes in
the base material. Thus every machining process also causes a change of the surface
rim zone vis-à-vis the base material. Whether these changes affect the functionality
of the workpiece must be tested and confirmed in individual cases. The distinction
is frequently made between the external and the internal boundary layer. Figure 2.35
shows the structure of both boundary layers.
The layers referred to here cannot be clearly defined; there are no fixed
boundaries between them.
For further reference, please refer to the following sources: [Schm36, Schl51,
Czic03].
2.5.1.1 The External Boundary Layer
The external boundary layer is located between the surrounding atmosphere and the
atoms of the base material embedded in the crystal lattice. It is generated via the
reaction between the material and the atmosphere during and after machining.
The external boundary layer encompasses the following individual layers:
32 2 Metrology and Workpiece Quality
1–10 nm
0.3–0.5 nm
10–100 µm
Oxide film
Adsorption layer
Grease-/ oil film 
External boundary layer
Internal boundary layer > 5 μm 
Matrix
Plastic deformed and hardened
Elastic und plastic deformed
Elastic deformed
Fig. 2.35 Layer structure of metallic surfaces, acc. to SCHMALTZ [Schm36]
Reaction Layers/Oxide Film
On the surface of metals which are either freshly machined or created through
breakage can form thin reaction layers/oxide films. These layers may only have
a thickness of 10–20 molecular layers. They may, for example, influence wetting
capability or adhesive behaviour.
Adsorption Layer
After a short duration of time of being exposed to the air, the oxide layer is covered
with an adsorbed layer of water and gas. This adsorption layer becomes considerably
important, for example, in the context of electrical resistance when metallic contacts
are present.
Grease/Oil Film
Coolants in particular cause a grease and oil film to deposit itself after machining on
the adsorption layer. This film ranges from 10 μm to several 100 μm in thickness.
A grease and oil film can still be detected even after cleansing. These films can
exert considerable influence interface formation in further galvanic, adhesion, or
PVD/CVD processing.
2.5.1.2 The Internal Boundary Layer
The internal boundary layer is the layer of the machined workpiece bordering on the
base material and thus possesses practically the same chemical composition as the
2.5 Inspection of the Workpiece Rim 33
Worn toolSharp-edged tool
Cutting tool material: HC
Material: C15
vc = 200 m/min
f = 0.1 mm
50 µm 50 µm
Fig. 2.36 Plastic deformation of the workpiece rim by drilling
base material. The physical structure and the expansion of this layer in the direction
of the workpiece interior depend both on the material used and the manufacturing
operation used. The transition from the affected rim structure to the unaffected base
structure is continuous.
Figure 2.36 shows a model of the plastic deformation of the structure of
case-hardened steel C15 after a drilling operation. Larger mechanical and ther-
mal stresses caused by worn tools lead to clear modifications of the textural
structure.
In steel material processing, for example, given high enough temperatures and
high cooling rates, a hardening zone may develop in the outermost surface layer. In
the layers below this, annealing processes are possible due to the lower temperatures
and cooling rates in these layers (Fig. 2.37). As shown in Fig. 2.37, a hardening zone
has developed which has an extremely negative effect on the flank bearing strength.
2.5.2 Inspection of the Surface Rim Zone
Since, in addition to surface geometry, the physical properties of the areas near
the surface play an essential role, these properties must be tested, too. Therefore
methods familiar in the field of materials testing are applied:
• material analysis
• structure and texture investigation
• hardness testing
• fracture mechanical testing
• residual stress measurement
Structural changes caused by increased thermal stresses of the rim zones near
the surface can be determined in structure and texture investigations which are
standardly used in materials science. This is done by metallographic preparations.
Crack detection in rim zones of workpieces can be either destructive or non-
destructive. A crack is a locally limited detachment of the material structure of small
width but often considerable length and depth. The crack can arise through internal
34 2 Metrology and Workpiece Quality
df
dwLF
RF
Left flank
Right flank
Grinding burn with new hardening zone
left flank
200 µmLF
Pitting at the tip of a tooth
200 µmRF
400
500
700
900
1000
300
800
600
Vi
ck
er
s 
ha
rd
ne
ss
 / 
HV
0,
05
0 2 4 6 8 10
Distance from edge d / mm 
Left flank
Fig. 2.37 Grinding burn during gear tooth grinding
tensions or through external force effects. For further reference, please refer to the
specified literature [Schu04].
2.5.2.1 Residual Stresses
Residual stresses are characteristic for the internal tension of a load-free workpiece.
This tension results from inhomogeneous elastic and elastic-plastic deformations
which are present without the influence of external forces and torques. For every
component, there is a balance of all the internal forces and torques created by resid-
ual stress. Residual stresses are subdivided into macro-residual stresses (type 1) and
micro-residual stresses (types 2 and 3). Residual stresses caused by mechanical and
thermal influences is treated in the chapters devoted to the relevant manufacturing
operations. The basic formation mechanisms of thermally and mechanically induced
residual stress are shown in Fig. 2.38.
Residual stresses become significant in the context of dynamic stresses. The
attempt is therefore often made to use manufacturing processes in finishing which
create compressive residual stresses in the surface rim zone in order tominimize
the probability of crack formation and crack growth. Such operations include, for
2.5 Inspection of the Workpiece Rim 35
Yielding point
Stra
in
Rel
eas
e
Deformation path
Te
ns
ile
(+)
Co
m
pr
es
sio
n(–
)
St
re
ss
Yielding point
Plastic deformation
sE–
Compressive residual
stress
(mechanical caused) 
Temperature
Cooling
Heating
High-temperature limit of elasticity
High-temperature limit of elasticity
Plastic deformation
Tensile residual
stress
(thermal caused) 
Te
ns
ile
 (+
)
Co
m
pr
es
sio
n 
(–)
St
re
ss
sE+
Fig. 2.38 Formation of residual stress
example, finish and surface rolling as well as shot peening. In contrast, tensile
residual stress is probable in electrical discharge machining, because of its ther-
mal operating principle; in grinding and all manufacturing operations which use
geometrically defined cutting edges, both mechanical and thermal load spectrums
occur during processing, a prediction of the stress state is not easily possible.
Methods of Analysis
For all processes, the metrological definition of residual stresses occur indirectly
through the measurement of strains, electromagnetic parameters or speed of sound.
Below, the basic methods used in manufacturing technology will be explained.
For further information on this topic, please refer to the specified sources [Peit92,
Hauk87].
Mechanical Methods
The mechanical methods of analysis include the borehole method and the toroidal
core method. These are destructive methods which release residual stresses in the
rim zone near the surface through the insertion of very small holes or ring grooves
in this zone, which causes an altered stress state and thus strains, which can be
measured with wire strain gauges, for example. The average residual surface tension
can then be calculated from the strains in consideration of the material behaviour.
By removing layers of the surface rim zone at different surface positions, dis-
tributions of residual stresses can be defined up to the depth corresponding to the
drill’s diameter. Since drilling or milling processes only cause reproducible strain
changes at a depth of several hundredths of millimetres, this method cannot be used
to determine residual surface stress with strong rim zone gradients.
36 2 Metrology and Workpiece Quality
X-ray and Neutron Diffraction
In research and practice methods are used which utilize the diffraction of monochro-
matic X- or neutron radiation on the crystalline lattice structure of the material
phases (Fig. 2.39).
The BRAGG condition applies here:
n · λ = 2 · a · sin � (2.3)
Altered lattice distances a of the material phases can be calculated very accurately
from the displacement of the diffraction lines, which in turn allow to draw con-
clusions about residual stresses. The chosen measuring direction defines the stress
components to be detected. This allows the definition of the entire stress tensor by
measuring in multiple directions. In order to calculate residual stresses from strain
measurements, the elastic properties of the material phases are used. In the case of
many materials, their anisotropic behaviour can even be taken into consideration
when measuring at the respective lattice levels. Figure 2.40 shows the customary
device for measuring residual stresses (goniometer).
Source of radiation
ϑ ϑ
ϑ
λ
a
Detector
Fig. 2.39 Principle of X-ray diffraction on crystalline lattice
X-ray tube
Detector
Swivel table Specimen
Rotating table
Fig. 2.40 Goniometer
2.5 Inspection of the Workpiece Rim 37
Diffraction methods demand high requirements on measurement technology and
on occupational safety (radiation protection). X-ray diffraction is used in practice to
define residual stresses. In research, neutron sources can also be accessed at selected
locations, with which it is possible to detect depth structures up to some millimeters
without removing surface layers.
The X-ray method has a depth of penetration which is limited to an extent
depending on the absorption properties of the material and the wavelength used
for the X-radiation. Given the usual measuring parameters, the average penetra-
tion depths vary between approximately 1 μm for tungsten carbide and 5 μm
for steel; significantly greater depths of penetration are possible for light metals.
Residual stress depth profiles with great rim depths can be produced through the
electrochemical removal of layers of the surface rim zone.
Magnetic Methods
Magnetic analysis methods use the influence of mechanical stresses on resetting
in ferromagnetic materials. These changes can be measured by means of an exter-
nally applied alternating magnetic field. The measurands are coercive field strength,
superposition permeability, dynamic magnetostriction and magnetic and acoustic
BARKHAUSEN noise. The simultaneous analysis of several magnetic parameters
improves the evaluation of the quantitative measuring data. Magnetic methods are
characterized by compact and easily manageable test set-ups. Depending on the
material and the range of frequencies analyzed, surface rim zones can be gauged
ranging from a few micrometers to over 1 mm. Disadvantegous is that these meth-
ods they are limited to ferromagnetic materials and have to be calibrated to known
material states.
Acoustic Methods
Residual stresses are generated through lattice strains of the material which result
in an altered propagation of structure-borne sound. This acoustic-elastic effect can
be used to analyze stresses by means of ultrasonic measurements. Such methods are
based on the assumption that the sound velocity is proportional to the lattice strain.
The applicational potential of these methods are limited above all by the material
state. This is because microstructural errors and textures exert as great an influence
on sound velocity as structural gradients.
Chapter 3
Fundamentals of Cutting
DIN 8580 defines machining as all process variants of the third main group
“Cutting”, in which form is altered by means of reducing material cohesion.
Deformation is achieved by means of a relative motion between the tool and the
workpiece that brings about a transfer of energy [DIN8580].
In the standard, this basic classification is further refined in order to categorise
machining procedures [DIN8589]. Machining is defined as follows: cutting, in
which layers of material are mechanically separated from a workpiece in the form
of chips by means of a cutting tool. According to DIN 8580, machining comprises
Groups 3.2 (machining with geometrically defined cutting edges) and 3.3 (machin-
ing with geometrically undefined cutting edges) in the manufacturing classification
system.
The first part of this compendium on manufacturing processes deals exclusively
with process variants of Group 3.2. For this reason, the term “machining” will only
be used in the sense of cutting with geometrically defined cutting edges. What all
processes in the group of geometrically defined cutting edges have in common is
that they use a tool, of which the cutting edge number, geometry and position to the
workpiece are determined.
Several concepts and terms are necessary to describe cutting part, which will be
described in the following.
3.1 The Cutting Part – Concepts and Terms
The concepts, designations and terms used to describe the geometry of the cutting
part are set down in DIN 6581. The cutting part is the active part of the tool where
the cutting wedges are located with the cutting edges. The idealized cutting wedge is
made up of two faces: A rake face and a flank face, which cut in a line, cutting edge
S. The angle between these two faces is designated as the wedge angle β. Figure 3.1
shows an idealized cutting wedge [DIN6581].
The rake face Aγ isthe face of the cutting edge where the chip runs off. The
flank face Aα is the face on the cutting wedge, which is turned towards the new
workpiece surface (the cut surface). These terms make it clear that the cutting wedge
39F. Klocke, Manufacturing Processes 1, RWTH edition,
DOI 10.1007/978-3-642-11979-8_3, C© Springer-Verlag Berlin Heidelberg 2011
40 3 Fundamentals of Cutting
Chip
Tool
Workpiece
Trace of plane
of flank face Aα
Trace of plane
of rake face Aγ
Cutting speed vc
Feed velocity vfEffective cutting speed ve
Trace
of plane
of cutsurface
β
η ϕ
Fig. 3.1 Description of the idealized cutting wedge
(tool) should always be regarded in connection with the workpiece. This means that
considerable importance should be attached to process kinematics.
One model concept is often used for the sake of a simplified description of pro-
cess kinematics, that of the selected cutting point. This model simplifies the actual
kinematics by summarizing the spatial velocity fields in one point, the selected cut-
ting point. At the selected cutting point, the velocity fields can be represented in a
summarizing fashion by means of vectors. These vectors can be summarized in turn
by vector addition in one total vector. Usually, the workpiece is assumed to be fixed;
all motions are carried out by the tool. The resulting velocity vector is designated
as the effective cutting speed ve. It can be divided into two components; the cutting
velocity vc in the cutting direction and the feed velocity vf in the feed direction. To
position the components of effective cutting speed clearly, two angles are defined:
• The effective cutting speed angle η as the angle between the effective cutting
direction and the direction of primary motion (see Fig. 3.1)
• The feed motion angle ϕ as the angle between the feed direction and the direction
of primary motion (see Fig. 3.1)
Since there are no ideally sharp tools in practice, cutting edge rounding is taken
into consideration (Fig. 3.2).
In almost all cases the transition between flank and rake face is curved. This
curvature is described by the cutting edge radius rß.
Up to this point, the concepts have been explained using a simple cutting wedge
formed by two faces. Generally, more complex tools are used, composed of several
cutting wedges, in the simplest case of one major cutting wedge and one minor
cutting face (Fig. 3.3).
3.1 The Cutting Part – Concepts and Terms 41
Tool
Trace of plane
of flank face Aα
Trace of plane of rake face Aγ
Chip
Workpiece
Cutting edge
 radius
vc
ve vf
rβ
β
ϕη
Fig. 3.2 Cutting edge with cutting edge radius
Tool shank
Direction of primary motion
Direction of feed
Major cutting edge S
Major flank face Aα
Minor flank face
Minor cutting edge S′
Rake face Aγ
Cutting edge
A'α
Fig. 3.3 Cutting edges and faces of the wedge, acc. to DIN 6581
Correspondingly, we speak of major and minor cutting edges. The major cutting
edge S is always turned towards the cut surface, the minor cutting edge S′ towards
the machined face [DIN6580]. If the selected cutting point is on the minor cutting
edge, the concepts defined in the former are named accordingly and furnished with
an apostrophe (′). The flat areas of the major and minor cutting edges are joined by
the corner radius r� (Fig. 3.6).
42 3 Fundamentals of Cutting
3.2 Reference Systems
In order to describe the location, position and direction of motion of a cutting wedge,
reference systems are utilized in which characteristic planes are defined that are
valid for all process variants.
The two standardized reference systems are the tool-in-hand system and the tool-
in-use system (Fig. 3.4). The tool-in-hand system was developed for tool design as
well as for the production and testing of cutting tools. In the tool-in-hand system,
the tool angle is measured without considering process kinematics. In actual cutting
processes however, the effective angle deviates from the nominal tool angle under
certain circumstances due to process kinematics (exception: the wedge angle). For
this reason, a tool-in-use system is also defined (Fig. 3.4, right).
The tool-in-hand system is a system, the reference plane of which is oriented
orthogonally to the assumed cutting direction. In contrast, the reference plane of
the tool-in-use system is oriented orthogonally to the effective direction. In both
systems, all planes contain the selected cutting point. The reference systems are in
agreement when the cutting direction corresponds to the effective direction.
First, the tool-in-hand system will be considered (Figs. 3.4, left, and 3.5):
• The basic plane, upon which all other planes are based, is the tool reference
plane Pr. It contains the rotation axis (if present) and lies perpendicularly to the
assumed cutting direction.
P
Pr
Pf vc
fv
v
Ppe
Pre
Pfe
Direction
of primary motion
Effective cutting
direction
Direction
of feed
Assumed direction
of primary motion
Assumed
direction of 
feed
ep
Fig. 3.4 Tool frame of reference (left) and tool-in-use system, acc. to DIN 6581
3.2 Reference Systems 43
Ps
Pr
Pn
Assumed direction
of primary motion
Assumed direction
of feed
Po
Fig. 3.5 Orientation in tool frame of reference
Tool
Trace of assumed
working plane Pf
Trace of tool cutting edge plane Ps
Trace of tool back plane Pp
Tool reference plane Pr
Assumed direction of feed
κr
re
Fig. 3.6 Position of tool cutting edge angle κ r
• The cutting edge plane Ps runs tangentially to the cutting edge S and perpendic-
ularly to the tool reference plane Pr.
• The tool orthogonal plane Po is perpendicular to the tool cutting edge plane Ps.
• The assumed working plane Pf is perpendicular to the tool reference plane Pr and
parallel to the assumed feed direction.
• The tool back plane Pp stands perpendicularly on the tool reference plane Pr and
perpendicularly on the assumed working plane Pf.
• The tool cutting edge normal plane Pn is perpendicular to the cutting edge S.
The tool cutting edge normal plane Pn is identical to the effective cutting edge
normal plane Pne, since it is not oriented to the tool reference plane but rather to
the major cutting edge.
The tool-in-use system is rotated towards the tool-in-hand system by the effective
cutting speed angle η. In tool-in-use system, the same signs are used as in tool-in-
hand system; they are followed however by an e, which stands for “effective”.
44 3 Fundamentals of Cutting
In order to designate all tool and effective angles clearly, the same index is
assigned to them that signifies the plane in which these angles are measured. For
example, the tool orthogonal wedge angle βo is measured in the tool orthogonal
plane Po, or the effective side rake angle γ fe is measured in the working plane Pfe.
Information about three angles around the three rotatory axes in space is required
in order to determine the orientation of the cutting edges or the positions in space
of the rake and flank faces. For an optimal cutting process, certain orientations are
responsible for three rotatory axes in space. In the following, these angles will be
explained exemplarily in the tool-in-hand system:
• the orientation angle around a rotation axis that stands orthogonally on the tool
reference plane Pr:
The tool cutting edge angle κ r between the tool cutting edge plane Ps and the
assumed working plane Pf, measured in the tool reference plane Pr, is defined
as the angle that determines the position of the major cutting edge S in the tool
reference plane Pr. It is measured mathematically positively on the basis of the
assumed working plane Pf (Fig. 3.6).
• the orientation angle around a rotation axis that stands orthogonally on the tool
cutting edge plane Ps:
The tool cuttingedge inclination λs between the major cutting edge S and the tool
reference plane Pr, measured in the tool cutting edge plane Ps, is defined as the
angle that determines the position of the major cutting edge S in the tool cutting
edge plane Ps (Fig. 3.7).
• the orientation angle around a rotation axis that stands orthogonally on the tool
cutting edge normal plane Pn:
This rotation is executed by one of the tool angles αn or γ n, which are mea-
sured in the tool cutting edge normal plane Pn. The tool normal rake angle γ n is
customarily used as the orientation angle, as it has a large effect on the cutting
process (Fig. 3.8).
Tool
Trace of
assumed
working plane Pf
Tool-cutting edge plane Ps
Major cutting edge S
Trace of
Tool reference
plane Pr
Assumed direction of
primary motion
λs
Fig. 3.7 Position of cutting edge inclination λs
3.3 Basic Process Variants 45
Tool
Trace of plane of flank 
face Aa
Trace of plane of rake face Ag
α
n
γn
βn
Trace of tool cutting
edge plane Ps
Cutting line of planes 
Pn and Pr
Assumed direction
of primary motion
Assumed direction
of feed
Tool cutting edge normal plane Pn
Fig. 3.8 Position of wedge in tool frame of reference
It is measured mathematically positively between the tool rake face Aγ and the
tool reference plane Pr in the tool cutting edge normal plane Pn. The tool normal
wedge angle βn is always fixed on the tool and cannot be changed by tool rota-
tion. The tool normal orthogonal clearance angle αn lies between the tool flank face
Aα and the tool cutting edge plane Ps, measured in the tool cutting edge normal
plane Pn.
The angle between the flank face Aα and the tool cutting edge plane Ps, measured
in the tool orthogonal plane Po, is designated as the tool orthogonal clearance angle
αo. The angle between the rake face Aγ and the tool reference plane Pr is defined as
the tool orthogonal rake angle γ o. The angle between the rake face Aγ and the flank
face Aα is called the tool orthogonal angle βo. As rule, the clearance angle, wedge
angle and rake angle – defined in the tool orthogonal plane Po, the assumed working
plane Pf and the tool cutting edge normal plane Pn – must in total make a 90◦ angle.
The kinematic factors of tool cutting edge angle κr, tool cutting edge inclination
λs and tool normal rake angle γ n are important influencing parameters on the cutting
process. For further information, see DIN 6582.
3.3 Basic Process Variants
Despite the large number of cutting process variants, they can still all be subdivided
into three main categories:
• free, orthogonal cuts,
• free, diagonal cuts and
• bound, diagonal cuts.
46 3 Fundamentals of Cutting
Workpiece
Tool
Chip
λ s
Workpiece
Tool
Chip
vc
vch
vch
vc
Fig. 3.9 Free, orthogonal and free, diagonal cut
The free, orthogonal cut is a special case, which can be brought about by the
following marginal conditions:
• Only the major cutting edge is being engaged (free).
• The tool cutting edge angle κ r is 90◦ (orthogonal).
• Tool cutting edge inclination λs is equal to 0◦ (orthogonal).
Practically speaking, this cut can be executed, for example, by means of longitu-
dinal face turning or cross cylindrical turning (see Sect. 9.1) allowing for the above
marginal conditions (Fig. 3.9).
The free, diagonal cut is more general and contains the free, orthogonal cut. The
following marginal conditions must be realized in the case of the free, diagonal cut:
• Only the major cutting edge is engaged (free).
• The tool cutting edge angle κ r can take on values that are not equal to 90◦
(diagonal).
• Arbitrary tool cutting edge inclinations λs are permissible (diagonal).
As soon as one of the last two marginal conditions is fulfilled, it is a free, diagonal
cut. Both conditions don’t have to be met simultaneously, although this can occur.
The general case, the bound, diagonal cut (Fig. 3.10), contains all abovemen-
tioned special cases and, in addition to the marginal conditions of the free, diagonal
cut, also permits the engagement of the minor cutting edge (bound).
All the basic process variants can be further extended by adding the following
categories:
• The uninterrupted cut and
• the interrupted cut.
The interrupted cut is the general variant among them, in which the cut takes
place only intermittently. In the case of the uninterrupted cut, temporal interruption
is infinitely small (the cut is continuous). Both process possibilities will be described
3.4 Chip Formation 47
Workpiece
Tool
Chip
vch
νc
Fig. 3.10 Bound diagonal cut
in great detail in later chapters (see Chaps. 9 and 10). In total, there are therefore six
basic process variants.
3.4 Chip Formation
At the beginning of the chip formation process, the cutting section penetrates the
material, causing it to deform elastically and plastically. After the maximum per-
missible material-dependent shear stress is exceeded, the material begins to flow.
Contingent on a given cutting section geometry, the deformed material forms a chip,
which runs off the rake face of the cutting section.
The property of plastic deformability is not solely related to the material; it can
also be brought about in a targeted way by altering the stress. The amount of stresses
is influenced by the feed velocity vf, the cutting speed vc and the depth of cut ap.
As far as process kinematics is concerned, the direction of cutting section stress is
determined by defining the tool normal rake angle γ n, the tool cutting edge angle κ r
and tool cutting edge inclination λs.
In order to ensure chip formation, a minimum chip thickness and depth of cut
must be exceeded (see Chap. 7 [Moll39, Djat52, Soko55, Bram61, Bram60]).
3.4.1 The Cutting Process
By varying the direction and amount of stress, either tough or brittle material
behaviour can be realized [Karm11, Böke14]. This has a large effect on chip for-
mation. During cutting, the direction of a particular stress can be set by the process
parameters tool normal rake angle, tool cutting edge angle and tool cutting edge
inclination. The amount of the particular stress is influenced by the cutting para-
meters cutting speed, feed velocity and depth of cut. To put it another way, the same
material can be tough or brittle depending on the direction and amount of stress,
48 3 Fundamentals of Cutting
Fig. 3.11 Model
representation of
deformation, acc. to Leopold
[Leop00]
independently of how it behaves at room temperature and under single-axis tensile
loads (tension test).
When cutting steel, one must bear in mind that brittle behaviour only occurs at
very low temperatures. This is not the case when machining cast iron, glass and
ceramics. There is brittle fracture behaviour for these materials even at high tem-
peratures due to their material structures. The particular stress must therefore be
optimally adjusted to the given material properties with process kinematics in both
direction and amount.
To clarify plastomechanical processes during chip formation, a grid is superim-
posed on the workpiece (Fig. 3.11) so that material deformation during cutting can
be observed (the visioplasticiy method) [Hast67, Chil71, Leop80, Leop00]. With
this method, the chip formation process can be made visible, and with the help of
this, chip formation models can be made. Finite element analysis is another possi-
bility for modelling chip formation. The visioplasticity method can help to verify
FE modellings, since it allows for a temporal and local definition of chip formation
on the superimposed grid [Leop80].
Special marginal conditions exist during chip formation: both high rake face tem-
peratures (about 770–1700◦C) as well as high deformation speeds (of an order of
magnitude of 104/s).
Figure 3.12 is a schematic representation of the chip formationprocess as
sketched with the help of a photograph of chip initiation (right).
In this representation, we can see a continuous plastic deformation that can be
subdivided into four zones. The transition from the workpiece structure (a) to the
chip structure (b) is made by simple shearing (shear zone). When cutting brit-
tle materials, minor deformation on the shear plane can already lead to material
detachment.
If however the material has higher deformability, detachment first occurs in front
of the cutting edge in zone (e). The tensile load under simultaneous perpendicularly
3.4 Chip Formation 49
Structure of the
workpiece
Cut surface
Tool
Flank face
rake face
Shear
plane
Structure
of the chip
c
vch
vc
a
Tool
Shear zone
0,1 mm
e
d
Cut surface
b
Fig. 3.12 Location of chip initiation
active pressure leads, together with the high temperatures prevalent here, to strong
deformations on the peripheries of the rake face (c) and cut surface (d). Sliding
over the tool surfaces causes further plastic deformations to arise in the boundary
layers. The “flow zone” (the non-etched white zone on the bottom of the chip), the
deformation texture of which forms parallel to the rake face, gives the impression
of a viscous flow process with an extremely high degree of deformation. The chip
resulting from the described chip formation process is designated as a continuous
chip. Other chip types include lamellar chips, segmented chips and discontinuous
chips.
Figure 3.13 shows a quantitative profile of normal and tangential stresses result-
ing from the resultant force components acting on the rake face. These stresses –
in conjunction with temperatures prevalent in the contact zone, which can amount
to over 1000◦C in the continuous chip formation zone – lead to deformations with
sheer strains ε¯ between 0.8 and 4.0 and sheer strain speeds ε¯˙ of up to 106/s. For the
sake of comparison, Fig. 3.13 provides corresponding figures from the tension test.
Cutting conditions under which cemented carbide tools operate result in deforma-
tion and material heating durations in the order of magnitude of milliseconds; the
heating velocities are theoretically around 106◦C/s [Opit70].
3.4.2 Different Types of Chip Formation
When chip formation has been ensured, it can occur in different ways. ERNST
[Erns38, Erns41] has made a phenomenological classification of chip formation
types. Continuous chip formation is characterized by an evenly deformed material
50 3 Fundamentals of Cutting
vch
d
Flow area
Tool
Detail A
vch
vc
T5
T4
T3
w
Distribution
of stress
Cutting temperature:
Chip flow velocity:
Heating velocity:
Average normal stress:
Average shear stress:
T1 ~ 1030 °C
vch = 67 m/min
106 °C/s
σn = 350 N/mm2
τ = 250 N/mm2
Shear strain
Shear strain rate
Tension test: ε ~ 0.2
Chipping: 0,8 < ε< 4.0
Tension test : ε~10–3 s–1
Chipping: ε ~ 104 s–1
Workpiece Material: C45E; Tool Material: HW-P20; ap = 2 mm; f = 0.25 mm; vc = 160 m/min
σn
σ
τ
τ
−
−
Fig. 3.13 Conditions of the cutting process, acc. to KÖNIG [Köni67]
chip structure, the cause of which is assumed to be in the temporally highly uniform
friction conditions between the chip and the tool.
Lamellar chip formation is characterized by an unevenly deformed material
structure between the chip and the tool, the cause of which is explained by tem-
porally highly altered friction conditions between the chip and the tool (stick-slip)
or by dynamic stress transfer [Scha64]. Locally enhanced structural deformations,
shear bands, can be recognized in the chip structure, that are characteristic of this
type of chip formation.
The frequencies of the vibrations caused by stick-slip are in the range of kilo-
hertz and have small amplitudes. The high level of local structural deformation is
3.4 Chip Formation 51
explained by the fact that the thermally caused material softening is more dominant
than its mechanical hardening. Research into these phenomena relies on the work of
ZENER and HOLLOMON, who have described them and conceived the term “shear
band”. Shear bands appear when machining high-strength materials with high lev-
els of deformability. They are also especially frequent when machining with high
cutting speeds [Hopp03].
If the stress condition in the shear zone exceeds the deformability of the mate-
rial (shear strength), there is a detachment of material areas, which then fuse with
each other again. This leads to the formation of segmented chips. This can be con-
ceived of as a special case of lamellar chip formation in which highly localized
deformations (shear bands) arise as well.
In addition to vibrations, the entire system – consisting of the machine tool,
tool, workpiece and fixtures – is influenced by further dynamic effects such as the
regenenerative effect or directional coupling [Weck77]. The frequencies of these
effects are much lower, in the area of a thousand hertz. In contrast to the higher-
frequency chip formation dynamics, these vibrations do not so much influence chip
formation in the sense of material deformation as much as chip formation from
outside, since they are the cause of dynamic alterations in the nominal feed.
SCHWERD has made microcinematographic experiments and suggested the con-
cept of “discontinuous chip” [Schw36]. This process basically distinguishes itself
from the other processes of chip formation by the fact that no plastic defor-
mation occurs before fracture, but rather fracture takes place without plastic
deformation. Discontinuous chips can be observed in the case of materials with
very brittle properties, e.g. cast iron, stone, fibre-reinforced plastic or titanium
alumnides.
Figure 3.14 summarizes the principal chip types [Vier70].
• Continuous chips form when the material has sufficient deformability (εB > ε0),
the microstructure is uniform in the cutting area, deformation does not cause
embrittlement and chip formation is not impaired by vibrations.
• Lamellar chips form when εB < ε0 < εF or the microstructure is not uniform or
vibrations lead to variations in chip thickness. Lamellar chips can form with high
feeds as well as with high cutting speeds.
• Segmented chips consist of chip segments that are separated in the shear plane
and fuse together again. The form when εF < ε0, whereby this is not only the
case for brittle materials like cast iron but also can come about if deformation
causes embrittlement in the microstructure. Segmented chips can also be formed
at extremely low cutting speeds (vc = 1 − 3 m/ min).
• Discontinuous chips most form when cutting brittle materials with uneven
microstructures such as certain types of cast iron and stone. The chips are
not detached, but are torn off the surface, often causing damage due to small
breakings from the workpiece surface.
Figure 3.15 gives a tool-oriented overview of the dynamic system of spheres of
influence, all complexly overlapping and mutually retroactive, which brings about
52 3 Fundamentals of Cutting
Continous chip1 Lamellar chip Segmented chip2 3 Discontious chip4
Range of lamellar-, 
segmented- and 
discontinous chip
Range of 
continous
chip
Sh
ea
r s
tre
ng
th
 τ
1
2
3
4
E
B
Elastic region
Plastic region
Yield range
Sh
ea
r s
tre
ng
th
 τ
Degree of deformation ε
ε 0
Strain ε
ε0: Degree of
 deformation
 in the shear plane
E: Elastic limit
B: Breaking limit
F: Fraction point
F
Fig. 3.14 Types of chip depending on material properties, acc. to VIEREGGE [Vier70]
Mechanism of chip formation:
• Continous chip formation
• Lamellar chip formation
• Segmented chip formation
• Discontionous chip formation
Range of temperature Material properties
Range of velocity
Tool geometry
Friction and Deformation
Toolwear
Process parameter
Fig. 3.15 Influence on chip formation
the various types of chip formation. This system is heavily influenced by the velocity
fields, the materials to be machined, tool geometry and process kinematics.
This section will not deal with low-frequency dynamic effects on the part of the
machine tool. See the relevant literature for further information [Kron54, Opit70,
Weck77, Alti00].
3.4 Chip Formation 53
Tool reference plane Pr
Quarz glass pane
Buffed and etched
Workpiece
Hollow spindle
Primary motion
Feed motion
Tool
Fig. 3.16 Microcinematograpy
One method with which chip formation is directly observable, is microcine-
matography, the first application of which on the cutting process was already carried
out by KURREIN and then later by KLOPSTOCK [Kurr05, Klop23, Klop26] (see
Fig. 3.16).
The development of this process has been advanced by several researchers in the
course of time [Schw36, Merc45, Oxle59, Spaa71, Warn74]. In this process, chip
formation is observed through a quartz glass pane while photos are taken or films
are recorded successively in short intervals. One disadvantage of this procedure is
the limitation to very low cutting speeds (vc = 1 cm/min). Nevertheless, the process
has helped to explicate basic phenomena.
Today there are high-speed cameras available that can capture 150,000 images
per second. With such cameras, the phenomena can be made visible at the lower
threshold at higher speeds or especially during micro-cutting.
In order to be able to study chip formation more effectively, various methods
have been developed for cut interruption. During the cutting process, chip roots
are held “in statu nascendi”, providing information about chip formation after sub-
sequent metallographic preparation. In the case of the equipment used today, the
tool is suddenly accelerated out of the cut. The acceleration is brought about by a
predetermined breaking point on the swivel device that breaks during the cutting
process due to an impulse caused by an explosion [Lola49] (Fig. 3.17). Gente has
recently developed a method in which the delay time and the masses moved can be
reduced in comparison with the “explosion method” [Gent02] (see also Fig. 3.17).
In this way, investigations into chip roots are also possible with higher cutting
speeds.
Other methods have been developed in which a chip root is produced in the
workpiece by means of a predetermined breaking point (brittle fracture, Fig. 3.18)
[Buda68]. The advantage is that that chip root is detached from the workpiece so that
the effective area can be spared during the following metallographic preparation.
54 3 Fundamentals of Cutting
Moment 1
Moment 2
Tool
Sample
Direction of
primary
motion
Gun
Moment 1
Moment 2
Baffle
Tool
Fig. 3.17 Interruption of cutting process, acc. to LOLADSE [Lola49] (left) and GENTE [Gent02]
(right)
Predetermined
breaking point
Buffed sheet steel
with high Young’s
modulus
Sa
m
ple
Tool
Fig. 3.18 Interruption of cutting process by predetermined breaking point, acc. to BUDA
[Buda68]
3.5 Kinematic Surface Roughness
Calculations of interpenetration between the tool and the workpiece can provide key
information about the cutting process, including the associated undeformed chip
cross-sections in the case of changing engagement parameters (see also gear man-
ufacture) or micro-geometrical surface parameters. The following shows by way
of example the generation of surface quality for a simple external turning process.
Since chip formation processes are ignored, we speak in this case of the generation
of kinematic surface roughness. For this, the penetration of the workpiece by the
tool is geometrically evaluated taking into consideration the kinematics in the tool
reference plane Pr (Fig. 3.19).
3.6 Mechanical and Thermal Strain on the Cutting Section 55
Area of validity: f < 2 · r cos(κ r + r – 90°) and κ r + r < 180°
z
x
M1 M2
r
r ε
r –Rt
f
f/2
4
f 2
r ε
2 −Rt −=
Rt
ε8 ⋅ r 
f 2=
or
κ r
Rtfε
ε 
ε ε ε
Fig. 3.19 Geometric ratio of engagement in the cutting process
3.6 Mechanical and Thermal Strain on the Cutting Section
The resultant force F, represented here in the turning process as an example, can
be decomposed into the components cutting force Fc, feed force Ff and passive
force Fp (Fig. 3.20). Currently, these resultant force components are usually detected
metrologically with the help of piezoelectric force sensors.
vc
Fc
Primary motion
(Workpiece)
Direction of feed
(Tool)
F : Resultant force
Fc : Cutting force
Ff : Feed force
Fp : Passive force
Fa : Active force
FD : Thrust force
Fz
Fa
FD
Fp
Ffvf
ve
vc : Cutting speed
vf : Feed velocity
ve : Effective cutting
 speed
Fig. 3.20 Resultant force and its components in the cutting process, acc. to DIN 6584
56 3 Fundamentals of Cutting
h
hch vch
ToolWorkpiece
Fc
vc Shear plane
Sectional view A-A
Fc Cutting force
Ff Feed force
Fp Passive force
Ffp Resultant of
Ff and Fp
FκN Normal force of
major cutting edge 
FκT Tangential force of
major cutting edge 
FγN Normal force of
rake face 
FγT Tangential force of
rake face 
FΦN Normal force of
shear plane
FΦT Tangential force of
shear plane 
FcγN Resultant of
FγT and FγN, or FΦT
and FΦN
Ff
FfpFp
vfhA
A
FκN
FκT
Fγ N
FΦN
FΦT
Fκ N
Fγ T
Fcγ N
h Undeformed chip thickness
hch Chip thickness
vc Cutting speed
vch Chip speed
αoe Tool orthogonal clearance
β 
β 
oe
Wedge angle
γ oe Tool orthogonal rake angle
κ re Working cutting edge angle
λ se Working cutting edge inclination
Shear angle
γoe
 oe
λse = 0°
κre
α oe
Φ
Φ
Fig. 3.21 Components of resultant force in working plane of reference (top) and working plane
(below)
The analysis of forces derived for the orthogonal cut by Merchant [Merc45,
Merc45a] is the basis for determining the forces acting on the tool cutting edge
(Fig. 3.21).
Assuming an ideally sharp cutting edge and neglecting flank face wear, the tan-
gential force Fγ T and the normal force Fγ N acting on the tool can be calculated from
the force components. The following applies as long as the inclination is λs = 0◦
and the influence of the minor cutting edge is slight:
Fγ N = Fc cos γo − (Ff sin κr + Fp cos κr) sin γo (3.1)
Fγ T = Fc sin γo + (Ff sin κr + Fp cos κr) cos γo (3.2)
3.6 Mechanical and Thermal Strain on the Cutting Section 57
where the bracket term in Eqs. (3.1) and (3.2) corresponds to the normal force of
the major cutting edge FκN (Fig. 3.21). Because of the (in most cases) small dif-
ference between the angles in the tool-in-hand system and the tool-in-use system,
calculations are made within the tool-in-hand system for the sake of simplicity.
For the case γo = 0◦, the cutting force Fc and the feed force Ff act perpen-
dicularly/tangentially to the rake face. In contrast to Merchant’s original work, a
neutral tool orthogonal rake angle (γo = 0◦) was selected, and the forces were
shown with the effective cutting direction on the tool cutting edge. Designation
of the acting forces and their signs is derived from the standards ISO 3002/4 and
DIN 6584.
As Fig. 3.22 clarifies for the simplest case of an orthogonal cross section with
a neutral tool orthogonal rake angle, the cutting and feed forces on the worn tool
are composed not only of the normal force of the rake face Fγ N and the tangential
force of the rake face Fγ T but also of the tangential force of the flank face FαT
and the normal force of the flank face FαN. The friction coefficient, which can be
calculated from the cut and feed forces, thus contains not only rake angle friction
but also forces acting on the flank face caused for example by elastic deformationof the workpiece surface and by tool wear [Klau65]. As research has shown, the
normal force of the flank face amounts to about 40% of the feed force and thus to
F
ΦN
F
a
= F
z
Fc
Ff
F Φ
T
Chip
Toolvf
vc
FγN
FγT
VB
FαT
FαN
Cutting
edge
acc to: Merchant
|Fc | = |FγN |+ |FαT |
|Ff | = |FγT |+ |FαN|
True for:
γo = 0°, κ r = 90°, VB > 0
h
Φ
Fig. 3.22 In the cutting
process the mechanical strain
of the cutting edge results
from normal and tangential
forces and their components
acting on rake and flank face,
acc. to MERCHANT [Merc45,
Merc45a]
58 3 Fundamentals of Cutting
about 66% of the rake face component in the feed direction [Spaa67, Lutz68]. A
percentage of approximately 10% of the cutting force Fc was ascertained for the
tangential force of the flank face.
The average normal and tangential stresses originating from the resultant force
components acting on the rake face are between 350 and 400 N/mm2 or between
250 and 350 N/mm2 when machining construction steel [Köni72].
Materials that are difficult to cut result in values of 1100 N/mm2. Their profile
is qualitatively reproduced in Fig. 3.13. The size and direction of the resultant force
are strongly influenced by the cutting parameters and cutting section geometries
used. Figure 3.23 shows the dependence of the static components of the resultant
cutting force Fc, Ff and Fp on the feed f, cutting speed vc, depth of cut ap and the
tool cutting edge angle κ r qualitatively in a linear coordinate system.
The extremes in the profiles of the resultant force components over cutting speed
can be ascribed to growth of built up-edge. The reduction of forces with increasing
cutting speed is caused by the reduction of material strength at higher temperatures.
The components of the resultant force increase proportionally over the depth of cut
ap. Yet this is only valid if the depth of cut is larger than the corner radius of the
tool. The profile of feed force Ff and passive force Fp over the tool cutting edge
angle κ r results from the geometric position of the cutting edge with respect to the
workpiece axis, since with a larger cutting edge angle the resultant force component
aimed in the feed direction increases, and its maximum is reached at κr = 90◦. If
the tool cutting edge angle is increased, the undeformed chip thickness h increases
proportionally to the reduction of the width of undeformed chip b. Since cutting
Cutting force Fc
Feed force Ff
Passive force Fp
Feed f
Cu
tti
ng
 fo
rc
e 
co
m
po
ne
nt
s
Fc
Ff
Fp
Cutting speed vc
Cu
tti
ng
 fo
rc
e 
co
m
po
ne
nt
s
Tool cutting edge angle κr
Fc
Ff
Fp
Cu
tti
ng
 fo
rc
e 
co
m
po
ne
nt
s
Depth of cut ap
Fc
Ff
Fp
Cu
tti
ng
 fo
rc
e 
co
m
po
ne
nt
s
Fig. 3.23 Components of
resultant force depending on
feed, cutting velocity, tool
cutting edge angle, and depth
of cut (qualitative)
3.6 Mechanical and Thermal Strain on the Cutting Section 59
Influencing variables
Change of cutting force components per degree
angle variation
Cutting force Fc Feed force Ff passive force Fp
Tool orthogonal
rake angle 
Tool cutting edge
inclination
Tool orthogonal
rake angle 
Tool cutting edge
inclination
D
ec
re
as
in
g
In
cr
e
a
si
ng
1.5 % 5.0 % 4.0 %
1.5 % 1.5 % 10.0 %
1.5 % 5.0 % 4.0 %
1.5 % 1.5 % 10.0 %
Fig. 3.24 Influence of cutting edge and rake angle on the components of the resultant force
force Fc is proportional over depth of cut ap(=ˆ width of undeformed chip b) but
increases degressively over feed (=ˆ undeformed chip thickness h), a light reduction
of Fc with increasing κr is the outcome of both changes.
Figure 3.24 provides some standard values for how the components of the resul-
tant force change if the tool orthogonal rake angle or the cutting edge inclination
is varied. This information can vary greatly however and should only be seen as
reference values.
Changing the tool orthogonal clearance angle in the range of 3 ≤ αo ≤ 12◦
has no mentionable effects on the resultant force components. Changing the corner
radius also does not have a significant effect on the forces as long as the condition
2r ≤ ap is met.
Tool wear (see Sect. 3.7) is another variable that influences resultant forces.
Varying effects on the components of the resultant force can be observed depending
on the type of wear.
Crater wear, resulting in a larger positive tool orthogonal rake angle, generally
leads of a reduction of cutting forces. On the other hand, these forces increase when
flank face wear is dominant since the friction surface between the workpiece and
the flank face becomes larger. Quantitative determinations concerning force increase
with increasing tool wear can only be approximated due to the large number of influ-
encing parameters. The following reference values can be assumed approximately
for force increase up to a width of flank wear land of VB = 0.5 mm: About 90% for
the feed force Ff, about 100% for the passive force Fp and about 20% for the cutting
force Fc.
Several specific characteristic values can be distinguished for calculating the
components of the resultant force. These component forces are directly proportional
to the width of undeformed chip b. To identify empirical laws, it is usually advis-
able to relate the measurement to already known, linearly dependent magnitudes in
60 3 Fundamentals of Cutting
Undeformed chip thickness
h / mm
mc
1
1 –mc = 0.7265
45°
1
0.1 0.2 0.4 0.6 0.8 1.0 2.0
200
400
600
800
1000
2000
Cu
tti
ng
 fo
rc
e 
F c
W
id
th
 o
f u
nd
ef
or
m
ed
 c
hi
p 
b
=
 F
c
’
/ (N
/m
m)
Fc' = kc1.1 ⋅h1–mc
kc1.1 = 1740 N/mm2
Fig. 3.25 Graphical determination of characteristic values kc1.1 and (1–mc)
order to minimize the number of variables going into the physical law. In this case,
we form the quotient Fc′ from the cutting force Fc and the width of undeformed
chip b. If we now plot the values thus found over undeformed chip thickness h in
a double logarithmic plot, the measurement points arrange themselves in a straight
line (Fig. 3.25).
The corresponding linear equation
log(Fc/b) = log(kc1.1) + (1 − mc) · log h (3.3)
can be converted into the KIENZLE Equation
Fc′ = kc1.1 · h(1−mc) (3.4)
The specific cutting force kc1.1 is the cutting force required to detach a chip of
undeformed chip width b = 1 mm and undeformed chip thickness h = 1 mm. The
exponent (1–mc) designates the gradient of the straight line Fc′ = f (h) in the double
logarithmic system.
To determine kc1.1 and (1–mc), cutting experiments are carried out for the combi-
nation of workpiece material and cutting tool material under investigation. In these
experiments, the relevant cutting forces are measured with constant cutting speed,
depth of cut and cutting section geometry and plotted in accordance with Fig. 3.25.
The required specific cutting force characteristic parameter kc1.1 is determined by
3.6 Mechanical and Thermal Strain on the Cutting Section 61
extrapolating the undeformed chip thickness to h = 1 mm. The tangent of the angle
between the straight line and the x-axis is the desired gradient value (1–mc).
Corresponding equations and characteristic values can be defined for resultant
force components Ff and Fp:
F’f = kf1.1 · h(1−mf) (3.5)
Fp′ = kp1.1 · h(1−mp) (3.6)
The values determined in this way are however only valid for undeformed chip
thicknesses of h > 0.1 mm. Values for kc and (1–m) can be found in [Köni82].
According to DIN 6584, the energy for cutting is the result of the product of the
paths to be travelled or paths travelled and the components of the resultant force act-
ing in their direction.Correspondingly the powers arising during the cutting process
result from the product of the speed components and the components of the result
force acting in their direction.
Cutting energy Wc and cutting power Pc:
Wc = lc · Fc (3.7)
Pc = vc · Fc (3.8)
Feed energy Wf and feed power Pf:
Wf = lf · Ff (3.9)
Pf = vf · Ff (3.10)
Effective energy We and effective power Pe are understood as the sum of all
corresponding cut and feed amounts:
We = Wc + Wf (3.11)
Pe = Pc + Pf (3.12)
Due to the relatively low feed speeds and feed paths, feed energy/power dur-
ing turning amounts to only about 0.03–3% of the corresponding cutting energy or
cutting power. For this reason, we can assume We ≈ Wc and Pe ≈ Pc for most
cases.
Figure 3.26 provides an overview of the segmentation of the total active energy
into shear, cutting and friction energy as a function of the undeformed chip thickness
[Vier70]. The illustration shows that the amounts of the different types of energy
depend on the undeformed chip thickness, whereby shear energy has the largest
share for most undeformed chip thicknesses.
The mechanical work used in cutting is almost completely converted into thermal
energy. Since the heat centres are identical with the deformation centres, the shear
62 3 Fundamentals of Cutting
Width of undeformed chip:
h / mm
Material: 55NiCrMoV6
Tensile strength: Rm = 800 N/mm2
Cutting speed: vc = 100 m/min
Width of undeformed chip:
b = 4.25 mm 
Tool orthogonal clearance:
αο = 5°
γο = 10°
Tool orthogonal rake angle:
0
200
400
600
1.0
100
300
500
700
0
Friction of flank face and cutting energy
Friction of rakeface
Total energy
Shear energy
Effective
energy We
We = Fe ·Ie
Deformation energy
Friction energy
Shear energy
Cutting energy
Friction of flank face
Friction of rake face
Latent energy
and
heat
Pa
th
 le
ng
th
 d
ue
 to
 p
rim
ar
y
m
o
tio
n 
Ef
fe
ct
iv
e 
en
er
gy
 W
e
/ (m
·da
N/
m)
0.2 0.4 0.6 0.8
Fig. 3.26 Segmentation of effective energy in cutting process depending on chip thickness, acc.
to VIEREGGE [Vier59, Vier70]
zone and the friction zones on the tool come into consideration as heat sources.
As Fig. 3.13 shows, the degree of deformation on the bottom of the chip is much
higher in the flow zone than in the shear zone, so one can expect the highest tem-
peratures between the chip and the tool. Since the thickness of the flow zone is very
small in comparison with the shear zone however, these higher temperatures should
not also be equated with a high energy conversion.
The illustration in Fig. 3.27 left gives information about heat that is
absorbed/dissipated by the workpiece, chip and tool. Most of the heat is dissipated
by the chip. Most of the mechanical energy (in this case 75% and generally more
than 50%) is converted in the shear zone. The heat arising in the individual develop-
ment locations is dissipated by thermal conduction, radiation and convection to the
environment. As a result of this heat balance, corresponding temperature fields form
in the workpiece and tool that change until equilibrium between added and removed
heat is achieved. The right side of the figure shows such a temperature field.
If we consider one material particle in the cutting zone, its temperautre will be
at least equal to that of a particle in the shear zone. As it glides further into the
contact zone, the material is intensively heated on the chip bottom side and the tool
on the rake face, because the energy required to overcome friction between the chip
3.6 Mechanical and Thermal Strain on the Cutting Section 63
600
700
650
600
400
450
500
300 310
380 oC
130
80
500
30
Workpiece
Chip
Tool
5%
2%
18%
75%
For steel
Material: Steel
Flow stress: kf = 850 N/mm2
Tool material: HW-P20
Cutting speed: vc = 60 m/min
Width of undeformed chip: h = 0.32 mm
Tool orthogonal rake angle: γo = 10o
Fig. 3.27 Distribution of heat and temperature in workpiece, chip, and tool in the process of steel
cutting, acc. to KRONENBERG [Kron54] and VIEREGGE [Vier59]
and the rake face is almost completely converted into heat. Since this process only
takes place in boundary layers of the chip and cutting tool material it heats up the
rake and chip bottom side all the more strongly the less time is available for heat
dissipation due to the higher cutting speeds. The maximum temperature does not
appear directly on the cutting edge, but rather at a certain distance from it on the
rake face depending on the cutting conditions.
Determining the temperatures arising during cutting has long been the subject
of intensive research [Gott25, Lang49, Schm53, Küst54, Vier55, Küst56, Axer55].
Already in 1956, KÜSTERS demonstrated temperatures of over 1000◦C on the rake
face when cutting steel with cemented carbide [Küst56]. Later investigations have
qualitatively confirmed these results again and again [Cass94, Denk90, Ehme70a,
Lenz66, Laus88, Beye72, Dama90, LoCa94]. The level of the temperatures acting
on the cutting section depends on the machined material, the cutting tool material,
the selected cutting conditions, tool wear and the cooling medium. Figure 3.28 pro-
vides an impression of the order of magnitude of the mean cutting temperatures to
be expected on the rake face as a function of cutting speed for various cutting tool
materials. In the vc = 20–50 m/min range, the temperature profile is non-linear in
the double logarithmic coordinate system. The reason for this is the the growth of
built up-edge arising in this cutting speed range (see Sect. 3.7.2), which interferes
with direct heat conduction.
In comparison to the rake face, temperatures on the flank face are generally low
(Fig. 3.29) [Denk90, Ehme70a, Axer55, Laus88, Beye72, Dama90]. The size of the
temperature difference between the rake face and the flank face, about 200–300◦C
64 3 Fundamentals of Cutting
Tool material: HW–P10, P30
High speed steel HS12–1–4–5
Material: C53E
Geometry of cutting section:
6° 6° 0° 70° 84° 0.8 mm
Cross section of undeformed chip:
ap · f = 3 x 0.25 mm2
Cutting time: tc = 15 s
300
350
400
450
500
550
600
650
700
750
800
850
900
950
1000
1050
1150
10 16 25 40 63 100 250
x
x
x
x
x
x
HW-P1
HW-P30
HS12–1–4–5
End of growth of–
build–up edge
Cutting speed vc
 
/ (m/min)
1100
Te
m
pe
ra
tu
re
 o
f r
ak
e 
fa
ce
 ϑ
 
/ °
C
α o γ o κ r ε r r ελs
Fig. 3.28 Mean rake face temperature
when cutting with cemented carbide tools according to [Ehme70a], is strongly co-
determined by the position of the contact zone on the rake face. The closer the rake
face moves towards the cutting edge, the more the temperatures prevalent on the
rake and flank faces conform. This means that temperatures are very high on the
flank face as well, especially during finishing operations with high cutting speeds
and small feeds. According to the information provided by the literature [Ehme70a,
Vier53, Schm53, Küst54, Vier55, Küst56, Axer55, Beye72], temperatures of over
800◦C can be found on the flank face depending on the selected marginal conditions
when cutting with cemented carbide tools.
There are various methods and devices available for temperature measurement
during the cutting process (Fig. 3.30) [Lowa67], which vary according to technique,
experimental setup and measurement position. The only metrological techniques
that come into question for measuring temperature during the cutting process are
time-resolution measurement techniques, since the cutting process duration (and
thus the achievement of a thermally stationary state) is too short for other methods
3.6Mechanical and Thermal Strain on the Cutting Section 65
0 1 2 3 4
1
2
1200°C
800°C
600°C
500°C
400 °C
Weld
Isolation of silicon
varnish 
CrNi-Wire
∅ 0.07 mm
mVCompensating
resistance 
Ni-tube
∅ 0.13−0.20
Example for the
distribution of
temperature
inside the
cutting edge 
Material: 30Mn4
Cutting tool material: HW-P10
Cutting speed: vc = 180 m/min
Cross section of
undeformed chip: ap·f = 3 ·0.74 mm2
Cutting wedge geometry
45° 90°
γo αo λs κγ ε r rε
Bevel width: 0.5 mm
45° to cutting edge0° 4° 0°
1000°C
Fig. 3.29 Measurement of temperature with embedded thermocouple on cemented carbide, acc.
to Küsters [Küst56]
2. Measurement with
 thermo-colours
1. Caloric heat
 measurement
3. Thermo-electric
 measurement
4. Radiation-
 measurement
2.1. Coat of paint at
 an uninvolved
 surface of the
 tool
2.2. Coat of paint
 at the inner
 surface of a
 splitted tool
4.1. Radiation
 pyrometer
3.1. One-chisel
 method
3.2. Two-chisel-
 method
3.3. Assembly of a
 complete
 thermocouple
3.4. Assembly of
 an isolated
 wire as
 themocouple
Temperature measurment during machining
4.3. Measurement
 of infrared
 radiation
 with PDS-cell
4.2. Infrared
 photo-
 thermometry
Fig. 3.30 Temperature measurement in cutting processes
66 3 Fundamentals of Cutting
(e.g. coating with temperature-sensitive colours or powders with constant melting
points). Of the processes introduced in Fig. 3.30, the one and two-chisel methods,
temperature measurement with thermocouples, pyrometers and photo thermometry
are of technical interest.
GOTTWEIN developed the one-chisel method as the first direct method for mea-
suring temperature and then, in cooperation with REICHEL, the two-chisel method
[Gott25, Herb26, Vier70]. Both methods are based upon the principle of a thermo-
couple. Both tool and workpiece form the heat soldering joints, while the tool clamp
is the cold soldering joint. The workpiece and the tool must be clamped in isola-
tion, since the thermoelectric voltage arising in the contact zone between the cutting
insert and the chip would otherwise be dismantled through the machine in a short
circuit. Contact between the passing chip and the workpiece outside of the contact
zone must therefore also be avoided. Both methods require electrically conductive
materials and cutting tool materials. In the case of the two-chisel method, cutting
tool materials with varying thermoelectric properties must be used. The greatest dis-
advantage of the single-chisel method is the protracted and elaborate process needed
to calibrate the thermocouple which must be carried out again for every cutting tool
material and workpiece material combination. This process can detect temperatures
of up to 1200◦C.
Temperature measurement with thermocouples is among the most common tech-
niques today. Installing a thermocouple into the tool or workpiece allows for a
point-wise determination of the temperature field (Fig. 3.29). The thermocouples
included are classified as encapsulated thermocouples, sheathed thermocouples and
single filament thermocouples. Thermocouples can be applied in blind holes in
the tool [Küst56, Qure66, Abra97] or in the workpiece [Osul02]. Temporal res-
olution is influenced by the response time of the thermocouple and heat transfer
between the thermocouple and the device under test. These techniques generally
have low temporal resolution. Also problematic is the contact heat transfer resis-
tance between the surface under test and the thermocouple due to the roughness of
the bore. This causes a difference in temperature between the measurement surface
and the thermocouple. In the case of sheathed thermocouples with isolated mea-
suring points, there is also the distance between the thermocouple surface and the
internal measurement point. Due to the extremely high temperature gradients with
short test times characteristic of cutting processes, this can lead to much lower mea-
surements. The maximum values dominant in the friction zones cannot be measured
in this way. Thermal compounds are used to improve heat transfer between the ther-
mocouple and the surface. Another disadvantage is that direct contact between the
thermocouple and the test object is necessary and that the holes used to position the
thermocouples can significantly affect the distribution of temperature.
The most important techniques in radiation measurement, which determines tem-
perature by measuring the heat radiation emitted from a surface, are pyrometry and
thermography. Pyrometry is the contact-free determination of absolute temperature
by measuring the inherent radiation of a body without spatial scanning of the object
field. Thermography provides a pictorial representation of temperature distribu-
tion. Radiation techniques have decisive advantages compared with thermoelectric
3.6 Mechanical and Thermal Strain on the Cutting Section 67
methods: the time resolution is much higher (whereby pyrometers are principally
faster than infrared cameras), and they are also contact-free.
One significant problem when measuring for an exact absolute temperature with
a radiation method is the dependence of the radiation emitted on the grade of
emission of the surface. Since the emission grade is the function of many factors like
temperature, wavelength, angular position, material and surface condition, calibrat-
ing the measurement device for a particular surface is very difficult. The precision of
total radiation and broadband partial radiation pyrometers are especially influenced
by factors that alter the spectral grade of emission of the surface. In cutting, effects
such as surface roughness and oxidation influence the grade of emission of different
surfaces greatly. To limit the influence of the grade of emission on measured tem-
peratures, narrow-band partial radiation, two-colour and multi-colour pyrometers
have been developed. The two-colour pyrometer (Fig. 3.31) has the advantage that
the spectral grades of emission ε1 and ε2 of the surface need not be known. Since
the two selected wavelengths lie directly next to each other, ελ1 ∼ ελ2. An error in
measurement will only result if both wavelengths λ1 and λ2 differ greatly. Further
advantages of this principle are that the measured temperature is independent of
signal dampening, due to dust for example, so long as both signals are dampened
equally. Moreover, the temperature of objects that are smaller than the optical field
of vision can be measured without error [Müll85].
Thermography with infrared cameras can be used as an alternative to pyrometri-
cal measurements. The advantage to this is the pictoral representation of temperature
information. Commercial cameras usually work with long wavelengths and large
broad-bands, which makes it possible to measure lower temperatures but which also
has a negative effect on the attainable accuracy. Scanning cameras that work with
a single detector are too slow for fast processes. Some high-speed infrared cameras
Insert
Workpiece Chip 
Point of measurement 
Lens
Dichroic
beam-splitter
Photodiode Amplifier
IR-filter
Glass fiber
(∅ = 0.42 mm)
Characteristics
• Measurement of two discrete wavelength bands 1 = 1.7 µm, and 2 = 2 µm
 which are set into ratio
• High temporal resolution (approx. 2 µs)
• Independence from grade of emission (no calibration required)
• Large range of temperature measurement (250–1200 °C)
• Contact free measurement 
λ λ
Fig. 3.31 Build-up of a two-colour pyrometer in principle, acc. to MÜLLER [Müll85]
68 3 Fundamentals of Cutting
offer integration times in the range of microseconds, making it possible to capture
fast processes without motion blur.
One major problem in the metrological determination of cutting tempera-
tures, besides poor accessibility, is the limitedlocal resolution of the available
measurement techniques as well as the small range (only a few micrometers) within
which the maximum temperatures arise. Local temperatures in the boundary layer
between the cutting tool material and the sliding chip also cannot be captured with
existing measurement methods. Promising attempts to determine thermal tool load
more accurately than previously have been made with sensors in thin-film technol-
ogy, which are applied to tool contact zones and determine temperature distribution
within the contact zones during the machining process [Zieb95, Töns95, Kloc97].
This technology is however not yet suitable for cutting processes with high wear.
Another highly promising method for determining temperatures in the boundary
layers between the material and the cutting tool material is the implantation of a
window made of diamond in the tool contact zone. Heat radiation emitted from the
chip bottom side passes through the window via a mirror integrated into the chisel
holder and reaches an infrared camera [Müll96].
A direct comparison of the temperatures measured by various authors is only
rarely possible, since experimental conditions generally deviate from each other.
There are numerous parameters that may have an effect on temperature: the type
of cutting process, materials, cutting tool materials, possible coatings of the cutting
tool materials, the use of cutting fluids and cutting parameters such as feed, cut-
ting speed, depth of cut and cutting section geometry. Moreover, the measurement
position is also crucial as well as metrological parameters like local or temporal res-
olution. The duration of the cutting process and the measurement time must also be
taken into consideration.
Cutting speed clearly has the greatest influence on the maximum temperatures
that arise in the contact zone between the chip and the tool. In the case of the
basic process, most measurements indicate an apparent rise in temperature at low
to average cutting speeds, while at high speeds there is an approximately constant
temperature level.
3.6.1 Influence of the Geometry of the Cutting Section
Depending on the cutting task at hand, we can select from a diverse array of cutting
section geometries. Geometry selection depends on
• the cutting tool material,
• material,
• cutting parameters and
• tool geometry.
Typical tool angles in steel cutting are given in Fig. 3.32. Determining the tool
angle always involves a compromise that can only do approximate justice to various
requirements.
Figure 3.33 shows how a change in cutting section geometry influences cutting
parameters.
3.6 Mechanical and Thermal Strain on the Cutting Section 69
Geometry of
cutting section
Cutting tool
material
Tool
orthogonal
 rake angle
Tool
orthogonal
clearance
Cutting
edge
inclination
λs κ r ε r
Cutting
edge
angle
Tool
included
angle
Corner
radius
rε
High speed steel
(HSS)
Cemented carbide
–6° to +20° 
–6° to +15°
–6° 
to 
+6°
10°
to
100°
60°
to 
120°
0.4 
to 
2 mm
6° to 8°
6° to 12°
αογο
Fig. 3.32 Tool angle for machining of steel
Lower minimum chip thickness Increasing stability of cutting section 
Improved chip formation
Improved surface
Reduced cutting force
Decreasing wear 
Decreasing 
wear 
Increasing stability
of cutting section 
Increased stability of cutting
section
Chamfer of the
rake face
Decreasing wear
Decreasing chatter,
Decreasing cutting
force 
 o = –10° to +20°
αo = 6° to 12°
κ r = 10° to 100°
rε = 0.4 to 2 mm
Decreasing chatter,
Decreasing cutting
force 
Increasing
surface quality Controlled
chip flow 
Increasing of passive forces,
Increasing stability of cutting
section 
 s = –6° to +6°
+
–
+–
γ
λ
Fig. 3.33 Influence of cutting edge geometry on characteristic machining values
3.6.1.1 Tool Orthogonal Clearance αo
Wear on the flank face (denoted by the width of flank wear land VB) is mostly
determined by the size of the tool orthogonal clearance. If it is large, the cutting
section is weakened in two respects: a heat build-up can ensue in the tool that can
possibly lead to impaired or loss of hot hardness; also, if the wedge angle is too
small, this increases the danger of breaking on the cutting edge. If αo → 0◦, surface
70 3 Fundamentals of Cutting
wear increases because pressure weldings appear on the contact areas of the friction
partners to a greater extent.
3.6.1.2 Tool Orthogonal Rake Angle γ o, Wedge Angle βo
The tool orthogonal rake angle γo can, in contrast to αo, be either positive or neg-
ative. It is responsible for the detachment of the material to be cut. The size of the
tool orthogonal rake angle γo affects the stability of the wedge; highly positive tool
orthogonal rake angles can thus lead to tool breakage as a result of cutting section
weakening. The primary advantages of a positive tool orthogonal rake angle are the
low cutting and feed forces as well as a usually improved workpiece surface qual-
ity. However, a chip flow that is supported by a positive tool orthogonal rake angle
often has insufficient chip breakage (tendency towards the formation of continuous
chips). Negative tool orthogonal rake angles increase cutting edge stability (used for
example for planing and machining workpieces with apertures, rolling skin or cast
skin). The thereby increased deformation of the passing chip and the large cutting
forces result in a high temperature load on the cutting section. Increased crater wear
appears on the rake face, which can lead to a decreased tool service life. Together
with the wedge angle βo, the tool orthogonal rake angle γo and the tool orthogonal
clearance form a right angle αo (Fig. 3.8).
3.6.1.3 Tool Included Angle εr
Due to the tool stability that is desired under extreme cutting conditions, the tool
included angle εr should be as large as possible. Small tool included angles are
required especially for copier and NC machining. The possible range is limited by
specifying the position of the major cutting edge and by making sure that the angle
between the minor cutting edge and the feed direction is at least 2◦ in order to avoid
subsequent shaving of the minor cutting edge on the workpiece.
3.6.1.4 Cutting Edge Angle κr
At constant feed and depth of cut, the width of undeformed chip b increases with
decreasing κ r. This causes the specific cutting edge wear to sink, so small cutting
edge angles are used especially to machine high strength materials in order to keep
tool load and wear low. On the other hand, the passive force Fp rises with decreasing
κ r, increasing the danger of clattering vibrations due to growing instability of the
cutting process.
3.6.1.5 Cutting Edge Inclination λs
A negative cutting edge inclination can stabilize the cutting process to a large extent
since the lead of the tool does not take place on the cutting edge but rather towards
the middle of the cutting edge. This results in an improved load profile, so that the
danger of cutting edge breakage due to local overloading is reduced. First cuts that
have minimal loads are of particular importance especially in the case of interrupted
3.7 Wear 71
cross sections (e.g. in milling or planing) as well as in the machining of cast iron
and forged parts (workpieces with transverse drill holes, shrinkage cavities).
Negative cutting edge inclinations induce large passive forces, which must be
absorbed by the machine tools (stiffness perpendicular to the main spindle!).
The cutting edge inclination also influences the direction of chip flow. A nega-
tive cutting edge inclination can result in the chip getting diverted to the workpiece
surface, decreasing the surface quality.
3.6.1.6 Corner Radius rε
The cornerradius to be selected depends on the feed f and the depth of cut
ap. Together with the selected feed, it influences the attainable workpiece sur-
face quality to a great extent (see Sect. 7.2.3), whereby the following relation is
approximately true:
Rt = f
2
8rε
(3.13)
Large corner radii improve the surface quality and cutting stability. Small corner
radii have the advantage of a smaller clattering tendency due to smaller passive
forces.
3.7 Wear
During the cutting process, deformation, separation and friction processes take place
in the area of the cutting edge. The cutting tool materials used are subject to an
extremely complex load collective characterized by high compressive stresses, high
cutting speeds and high temperatures.
Using cutting parameters common in practice, cutting tools reach the end of their
service life because of continuously increasing wear on both rake and flank faces.
This is explained as the progressive loss of material from the surface of a solid
body, brought about by mechanical causes, i.e. contact and relative motion of a
solid, liquid or gaseous counter body.
3.7.1 Wear Mechanisms
The main mechanisms that cause wear are adhesion, abrasion, tribochemical
reactions and surface disruption (Fig. 3.34).
3.7.1.1 Abrasion
Abrasion occurs when rough areas of the counter body or particles present as inter-
mediate material or also as counter bodies penetrate into the surface of the cutting
tool material and simultaneously make a tangential motion, producing scores and
72 3 Fundamentals of Cutting
Adhesion Abrasion Tribochemical
reaction
Surface damage
Fig. 3.34 Schematic illustration of the four main mechanisms of wear, acc. to ZUM GAHR
[ZumG92]
Micro-plowing Micro-cutting Microcracks
Fig. 3.35 Types of material damages caused by abrasive parts, acc. to ZUM GAHR [ZumG87]
micro-cutting. The various forms of abrasion are designated as groove, fluid ero-
sion, mill, notch or blast wear. Often, the term “groove wear” is used synonymously
with the term “abrasive wear” [Habi80, Czic06].
In the case of groove wear, hard areas of roughness or hard particles penetrate
into the surface of the stressed material, creating scores or grooves by a sliding
movement. Wear can be caused by a rough, compact counter body (counter body
grooving) or by loose particles in the case of sliding load (particle grooving).
The process of material damage can be subdivided into micro-plowing, micro-
cutting and microcracks (Fig. 3.35). Micro-plowing and micro-cutting are the
predominant wear processes in the case of ductile materials. In the case of micro-
plowing, the material is plastically deformed within the wear groove and pushed
towards the groove edges. In the ideal case, no material is removed. This can how-
ever occur if the material is repeatedly pushed to the groove edges due to the
simultaneous effect of many abrasive particles or due to multiple effect of a sin-
gle particle, and the material finally fails due to fatigue. Ideal micro-cutting leads to
material removal in the form of a chip, the volume of which is equal to the volume
of the wear groove created. In the case of brittle materials, microcracks also arise.
Material particles are hereby generated by means of cracking and crack growth in
the stressed surface. The volume of the wear particle created in micro-cracking is
much larger than that of the continuous wear groove [ZumG87, ZumG92, Habi80].
3.7 Wear 73
3.7.1.2 Adhesion
Adhesion is defined as the formation of bonds between certain molecules [Erin90,
ZumG87]. As a wear mechanism this process is understood as one in which atomic
bonds (e.g. in the form of micro-welds) are formed in the contact zone between the
material and the cutting tool material. These bonds are then deformed, reinforced
and sheared off during a tangential motion of the friction partners.
The process of adhesion (i.e. the formation of micro-welds between two fric-
tion partners) can come about by atomic interaction (chemical adhesion) between
the partners, such as thermally induced diffusion processes, electron exchange or
electric polarisation. It can also result from mechanical snagging (mechanical adhe-
sion) of the workpiece material, which has plastically become extremely deformable
under high temperatures, with the cutting tool material [Tell93].
By means of adhesion, particles can be transferred from one wear partner to the
other. If this transfer is the primary cause of wear, it is sometimes referred to as
adhesive wear, although other wear mechanisms are usually also involved in the
formation of loose wear particles [Habi80].
Adhesion is a pairing property that depends on the characteristics of the base
and counter bodies. Whether a material has the propensity to adhesion can only
be answered with reference to the material of the counter body, with which it may
either form strong, weak or no adhesion bonds at all [Habi80].
The strength of an adhesion bond is described by the adhesion coefficient. This
is defined as the quotient of the normal force FN with which two solid bodies in
relative motion are pressed against each other and the opposing force FA that must
be applied to undo the bond formed by adhesion [Habi80, Andr59, Siko63].
The propensity of solid bodies to form adhesion bonds can be evaluated with
the help of their surface energy. Contact angle measurement is one method for
ascertaining the surface energy of solid bodies. [Bobz00] describes a measurement
process that makes it possible to compare the surface energies of coating systems
with respect to their polar and dispersive parts. Since it is the polar energy share that
is primarily responsible for the propensity of a solid body to adhesion, the surface
energy of cutting tools (i.e. especially that of the coating system) should have as little
polar energy as possible in order to minimize that propensity [Bobz00, Kloc05].
During chip formation, new material surfaces are generated that, without adsorp-
tion or reaction boundary layers in statu nascendi, come in a chemically highly
active state into contact with the tool surface under high pressures and temperatures,
approaching that surface on the atomic level. Plastic deformation processes lead to
the formation of large real contact areas. The cutting process thus provides very
good opportunity for the appearance of adhesively caused interactions [Bömc89,
Neis94, Erin90].
3.7.1.3 Tribooxidation
The term “tribooxidation” refers to chemical reactions of the cutting material and the
material with components of the intermediate material or the surrounding medium
as a result of activation caused by friction. Tribooxidation changes the properties
74 3 Fundamentals of Cutting
of the external boundary layer. Reaction products can be formed that are either
removed with the chip or remain stuck to the cutting tool material as a coating.
Wear can be increased or reduced by this. Whether tribooxidation increases the
amount of wear essentially depends on the hardness of the reaction products formed
in comparison with the hardness of the cutting material. A decrease is especially
possible when the reaction layers prevent direct metallic contact between the base
and counter bodies, limiting the effect of adhesion [Erin90, Habi80, Tell93].
3.7.1.4 Diffusion
Diffusion is the thermally activated change of position of individual atoms
[Horn67]. This involves a temperature-dependent physicochemical process in which
the wear resistance of the cutting tool material can be reduced by foreign substances
diffusing with it or its own components diffusing away from it. The diffusion of
atoms from the tribologically loaded areas of the tribopartners leads on the one hand
to a direct loss of material, which is usually very small but can still certainly be mea-
sured as a quantity of wear; however, much more serious is the potentialreduction
of the wear resistance of cutting tools by the diffusion of certain alloying elements.
Diffusion of essential alloying elements can lead to decreased hardness and thus to
reduced resistance of the cutting tool material to abrasion [Neis94, Erin90, Habi80].
The high pressures and temperatures in the contact zones characteristic of the
cutting process provide an ideal setting for diffusion processes between the material
and cutting tool material. Diffusion processes appear especially when high cutting
speeds are used, thereby leading to high temperatures in the contact zone. In the
area of the contact zone, the cutting tool material and the material approach each
other on the atomic level. Furthermore, there is a difference in concentration due to
the differing compositions of the material and the cutting tool material which also
is preserved in the material because of the constantly added chip [Neis94, Erin90].
3.7.1.5 Surface Damage
Surface damage occurs as a result of tribological alternating stresses. In the stresses
surface areas, alternating mechanical stresses lead to structural changes, fatigue,
cracking, crack growth and even to separation of wear particles [Habi80].
As opposed to abrasion, in which wear particles can be formed by a single stress
process, surface damage is usually preceded by a longer incubation period during
which no measurable wear occurs. In this period, the formation of wear particles
by means of structural changes as well as cranking and crack growth is prepared
[Habi80].
3.7.2 Causes of Wear
Friction processes in tool contact zones are comparable to those of dry friction in
a vacuum. Together with extraordinarily high mechanical and thermal stresses, the
tool is generally worn out quickly [Rabi65, Krag71, Opit70a, Opit70b].
3.7 Wear 75
To
ta
l w
ea
r
Cutting temperature
(cutting velocity; feed et al.)
Adhesion
High temp.
oxidation 
Diffusive
wear
Abrasion
Fig. 3.36 Causes of wear in cutting processes
According to current information, the following separate causes can be given for
the collective term “wear” (Fig. 3.36):
• mechanical wear (abrasion),
• the shearing off of adhered material
• damage to the cutting edge due to mechanical and thermal overstress,
• diffusion,
• scaling.
These processes overlap to a large extent and are only partially separable from
each other with respect both to their cause and to their effect on wear [Opit67,
Köni65, Ehme70b, Ehme70c].
3.7.2.1 Mechanical Wear (Abrasion)
Mechanical wear or abrasive wear occurs at both low and high cutting speeds.
Groove wear in the form of counter body or particle grooving can be seen as the
dominant form of wear in this category. In the case of counter body grooving, tri-
bological stress on the cutting tool material is based on the abrasive effect of hard
particles fixed on the contact surface of the workpiece or the chip. These particles
can originate from the workpiece material (oxides, carbides, nitrides) or be trans-
ferred by adhesion from the cutting tool material to the workpiece or chip bottom
side. In the case of particle grooving, loose particles are the cause of wear. These
can arise directly from abrasion (micro-cutting, microcracks) or from surface dis-
ruption. However, these may also be adhesion particles or products of tribooxidation
removed by abrasion or surface fatigue. Due to the high pressures and temperatures
that predominate in the contact zones on both rake and flank faces, we must presume
that the loose particles generated are pressed into the softer counter bodies flowing
by them, contributing to further wear by counter body grooving.
76 3 Fundamentals of Cutting
Due to abrasion or surface disruption on the cutting edge or in the contact zones
on the rake and flank faces, cutting tool material particles that have broken away
from the cutting tool material flow over the rake or flank face under high pressure.
This can cause further wear by micro-cutting or microcracks. This process, also
called “self-wear” [Ehme70a], is of particular importance especially with respect to
the formation and development of flank face wear.
3.7.2.2 Shearing-Off of Adhered Material Particles (Adhesion)
In the case of micro-welds being sheared off, material separation can occur in the
boundary layer, within one or within both bodies. The term adhesive wear is used as
soon as material is separated in the cutting tool material. Adhesion is also responsi-
ble for the formation and growth of built-up edges, in which material is transferred
from the material to be cut to the cutting tool [Erin90, Habi80, ZumG87].
Ferritic and austenitic steel materials have a high propensity to adhesion with the
cutting tool material. The reason for this is above all the high plastic deformability
of these materials. The high ductility of ferritic materials is based above all on their
relatively low strength, in the case of austenitic steel materials on their face-centred
cubic crystal lattice.
Tungsten carbide, the basis of hardness and wear resistance in conventional WC-
Co cemented carbides, has a hexagonal crystal structure. On the other hand, the
crystal lattice of the binding metal cobalt is face-centred cubic above 690 K, which
is favourable for adhesive processes. The titanium-based coating systems commonly
deposited on cemented carbides also have face-centred cubic structures, resulting in
a strong tendency to adhesion when machining austenitic steels. The types of wear
resulting from this strong adhesive tendency can range from material bonding on
the rake and flank faces to de-coating of coated tools in the area of the contact zone.
Built-up edges are highly reinforced layers of the machined material that take
over the function of the tool cutting edge as bondings on the tool. This is made pos-
sible by the property of certain materials to harden during plastic deformation. The
material adhering to the cutting edge is deformed by chip pressure, making it very
hard. This makes it possible for it to take over the function of a chip-removing tool.
Depending on the cutting conditions, built-up edge particles slip periodically
between the flank face and the cutting surface. In the case of high hardness and
removal frequencies up to about 1.5 kHz, these particles lead to increased flank face
wear and considerably deteriorate the surface quality of the workpiece (Fig. 3.37).
Since the chip is diverted with the built-up edge and not the rake face, crater wear is
usually negligible.
Figure 3.38 shows a wear-cutting speed function (VB-vc graph). According to it,
flank face wear does not increase with cutting speed continuously, but has at least
two distinct extremes [Opit64]. Wear first reaches a maximum at the cutting speed at
which the built-up edges reach their largest dimensions. A wear minimum appears
at the cutting speed at which no more built-up edges form.
Flank face wear decreases after exceeding the maximum, despite higher cutting
speeds. This can be ascribed to the fact that reinforcement of the built-up edge
3.7 Wear 77
1 2
3 4
Fig. 3.37 Scheme of the periodic growth of build-up edge
W
id
th
 o
f f
la
nk
 w
ea
r l
an
d 
VB
 / 
m
m
Cutting speed vc / (m/min)
100
0,24
0,20
0,16
0,12
0,08
0,04
0
1 2 3 5 10 20 30 50
Material: C53E
Cutting mat.: HW-P30
Cross section of
undeformed chip:
ap· f = 2 · 0.315 mm2
–4°8° 10° 90° 60° 1mm 
re
Geometry of cutting wedge:
Cutting time tc = 30 min
vc = 30 m/min
vc = 20 m/min
αo γo ε rκ rλs
Fig. 3.38 Wear on flank face and growth of build-up edge
begins to decrease as a result of recrystallization processes. It becomes unstable
and no longer migrates partially between the cut surface and the flank face, but as a
whole across the rake face.
The position of the maxima and minima of the VB-vc graphis contingent on
temperature. It is shifted to lower cutting speeds by any kind of measures taken
78 3 Fundamentals of Cutting
to increase the cutting temperature (e.g. higher feed, smaller tool orthogonal rake
angle, higher material strength). Measures to reduce the cutting temperature (e.g.
cooling) shift the extremes accordingly to higher cutting speeds [Opit69, Peke74].
3.7.2.3 Mechanical and Thermal Overstress
Damage to the cutting edge, such as breaking, parallel cracks, comb cracks or plastic
deformation, is the consequence of mechanical or thermal overstress.
Cutting Edge Chipping
Large cutting forces lead easily to breaking on the cutting edge (cutting edge chip-
ping) or corner if the wedge angle or tool included angles of the tool are too small or
the cutting tool material used is too brittle. In the case of such breaking, the profile of
the fracture surface is determined by the direction of the cutting force [Köni75]. Cut
interruptions can also lead to breaking, especially when tough materials are being
cut, the chips of which tend to stick.
Small amounts of breaking occur when the workpieces contain hard non-metallic
inclusions that originate during deoxidation of the steel [Opit64a, Opit66, Opit62].
Sintered oxides and more wear-resistant types of cemented carbide are sensitive to
this type of local overstress, especially in the case of manufacturing processes with
relatively small cross-sections of undeformed chip (e.g. reaming or shaving).
Distinct breaking of cutting tool material on the major and/or minor cutting edge
can also be caused by chips striking against the cutting edge or during the rotation
of shaft sections because of chip jamming between the cutting insert and the work-
piece. Breakage of the cutting edge can occur both on the top and bottom sides of
the insert.
Parallel Cracks
In the case of interrupted cuts (e.g. milling), the cutting tool material is subject to
strong mechanical alternating stress. This dynamic pressure threshold load can lead
to fatigue failure. Quick consecutive cutting force changes lead, especially in the
case of milling with cemented carbide tools, to parallel cracks (Fig. 3.39).
The quickly alternating stress in the formation of lamellar chips can also lead
to parallel crack formation if a critical stress cycle is exceeded [Domk74, Beck69],
e.g. when cutting titanium materials.
Comb Cracks
Comb cracks are a form of damage to the cutting edge as a result of thermal
alternating stresses (Fig. 3.39). Such stresses originate mainly while working with
interrupted cuts.
During tool engagement, the cutting edge quickly heats up to high temperatures.
It cools down after the tool exits the workpiece. The difference between the highest
and the lowest temperature depends, among other things, on the material, cutting
3.7 Wear 79
Heating
during cutting process
–y
Temperature Stress
Tensile + 0 – Compressive
Temperature
Tensile + 0 – Compressive
Stress
Cooling
Directio
n of pa
rallel
cracks
Di
re
ct
io
n 
of
pr
im
ar
y 
m
ot
io
n
Direction of comb cracks
Rake
 face
Flank
 face
VB
KM
42CrMo4+QT
Comb and parallel cracks
Comb cracks
vc = 275 m/min 
GJS70 vc = 200 m/min 
Fig. 3.39 Comb and parallel crack formation in milling, acc. to VIEREGGE [Vier59]
conditions and the ratio of the paths covered in the material and in the air. In the case
of interrupted cuts, the use of cutting fluids is particularly important with respect
to the size of the temperature difference because it has a much larger quenching
effect than air. Cooling is advantageous to the formation of comb cracks in the case
of cemented carbides and ceramic cutting tool materials. The profile of the comb
cracks is aligned with that of the isotherms of the temperature field in the cutting
part.
The comb cracks arising in the cutting part during milling or also in short-cycle
turning are generally temperature change cracks. These should be distinguished
from thermoshock cracks such as develop during turning or milling while using
cutting fluid. Thermoshock cracks are formed by a nonrecurring steep temperature
change, while temperature change cracks form only gradually in the cutting part
in the course of the cutting edge engagement. While in the case of thermoshock
the thermally caused tensile strain arising in the contact zone exceeds the strength
of the cutting tool material and leads directly to cracking, the stresses leading to
the formation of temperature change cracks are below the strength of the cutting
tool material. The development of temperature change cracks thus requires several
periodic temperature changes as a rule.
80 3 Fundamentals of Cutting
Due to the high energy level thermoshock cracks propagate with greater speed.
Crack growth is largely transcrystalline. Depending on the stress level, they can
acquire very large depths in the cutting tool material (> 1 mm), thus leading to
considerable weakening of the cutting edge [Gers98].
Temperature change cracks can be observed in both wet and dry cuts. Besides the
amount of thermal load, the cutting cycle and the number of cutting cycles play a
decisive role in their formation. The use of cutting fluids leads to a strong reduction
of the amount of cutting cycles required for cracking [Gers98].
Depending on the bending stress which the cutting edge is subject to during
the cutting process and which are superimposed over thermally induced stresses,
the cracks become larger with increasing numbers of cutting cycles. With a corre-
sponding output length, they can relatively quickly achieve a critical size leading
to fracture. As a result of the notching effect coming from the crack, they can also
become the source of further cracks [Gers98].
Plastic Deformation
Plastic deformation of the cutting edge arises when the thermomechanical stress
effecting the tool cutting edge exceeds the deformation resistance of the cutting tool
material. The influencing parameters are the strength of the material to be cut, the
cutting parameters, the geometry of the insert and, on the part of the cutting tool
material, its high temperature properties like hardness, compressive strength and
creep behaviour.
Cutting edges made of tool steel or high speed steel deform the stronger the
smaller the difference is between the temperature of the cutting edge and the anneal-
ing temperature of the cutting tool material. Plastic deformations also appear in the
case of cemented carbides and cermet, however only at higher temperatures (cutting
speeds) and under higher forces as is the case for tool or high speed steels. Cemented
carbides deform more the more binder phase there is, usually cobalt.
Plastic deformation of the cutting edge results in a considerable increase in wear
and can lead to sudden disruption of the tool by cracking and shearing off of the cut-
ting edge. It thus limits above all the cross-section of undeformed chip and cutting
speeds applicable in rough turning.
3.7.2.4 Diffusion
In the case of heat wear resistant cemented carbide tools, diffusion wear must be
expected at high cutting speeds and mutual solubility of the partners. Tool steel and
high speed steel already become soft at temperatures at which diffusion can hardly
manifest itself (e.g. about 600◦C for high speed steel).
When cutting with uncoated cemented carbide substrates, the following reactions
can occur (Fig. 3.40):
• diffusion of Fe into the binder phase Co,
• diffusion of Co into the steel, whereby Fe and Co form a gapless series of mixed
crystals,
3.7 Wear 81
WC: 
Dissolution into: 
Fe3W3C; (FeW)6C;
(FeW)23C6
(Ti, Ta, Nb, W) C
Co-W-C-mixedcrystal
ChipTool
Fe
C
Co
X6CrNiMoTi17−12−2
0
10
20
30
40
50
M
as
s 
pe
rc
en
t %
55µm
W
CoFe, Cr, Ni
Fe
W, Co
0
MaterialCutting material
Cemented carbide WC-Co
Fig. 3.40 Presentation of the diffusion phenomena in cemented carbide tools
• dissolution of tungsten carbide to form mixed and double carbides in the form of
Fe3W3C, (FeW)6C and (FeW)23C6.
The carbon released during the dissolution of tungsten carbide migrates in the
direction of lesser concentration, i.e. into the steel. The diffusion of carbon passes
into the cobalt phase. The maximum solubility of carbon in cobalt is about 0.7% at
1200◦C. In the presence of Fe the solubility is increased to 1.5–2%. The diffusing
iron thus introduces two reactions, both of which accelerate dissolution. It lends
itself to the formation of iron mixed carbides and increases the reception of carbon
in cobalt, which is in turn the prerequisite for the dissolution of tungsten carbide.
The diffusion processes that are possible between cemented carbide and steel
materials can be shown very effectively with the help of annealing tests. To this pur-
pose, cemented carbide-material pairs were annealed at a temperature of 1100◦C,
a pressure of 12.6 MPa and a holding time of 2 h in a protective gas atmosphere
of argon. Then the samples that were fused together on the contact surface were
separated on an abrasive cut-off machine perpendicularly to the contact level, and
82 3 Fundamentals of Cutting
after these sample halves were grinded and buffed, the distribution of elements
was analysed with a microprobe perpendicularly to the contact surface cutting tool
material/material.
As the quantitative line scans in Fig. 3.40 show, diffusion phenomena take place
under the selected annealing parameters between the uncoated cemented carbide
and the austenitic steel material here selected. On the side of the steel material,
there is a diffusion of iron, chrome and nickel into the cutting tool material. In the
opposite direction (i.e. from the cutting material into the material) tungsten and
cobalt diffuse.
The dissolution of tungsten carbide, the bearer of hardness and wear resistance,
caused by the diffusion of iron into the cemented carbide leads to a weakening
of the structure and lowers the resistance of the cemented carbide substrate against
abrasion. The consequence of this is the formation of distinct crater wear (Fig. 3.41).
Since diffusion phenomena depend on temperature, the cutting speed influences on
an elementary level the speed with which crater wear develops. Uncoated cemented
carbides can thus only be used at relatively low cutting speeds (vc < 100 m/ min).
The cemented carbide substrate also has a serious effect on crater formation. Mixed
carbides on the basis of titanium and tantalum/niobium have, in comparison to WC,
a much higher chemical stability towards iron. Mixed carbides reduce the diffusion
of iron into the cobalt binder phase, thus increasing the resistance of the cutting tool
material to crater wear and increasing its heat wear resistance quite substantially.
The most effective way to reduce crater wear is coating the cemented carbide
substrate. Cemented carbide layers seal the substrate surface. This prevents direct
contact between the substrate and the material, thus reducing or suppressing diffu-
sion phenomena. As the line scans of annealing tests in Fig. 3.42 show, diffusion
processes between the cemented carbide substrate and the material can by all means
still occur in the case of TiN or TiCN layers. In contrast, no more diffusion takes
place in the case of the coated cemented carbide with Al2O3 as an intermediate layer.
The aluminium oxide layer thus functions as a diffusion barrier. The elements iron,
nickel and chrome are found in the ZrN surface layer, while tungsten and cobalt are
in the TiN boundary layer (Fig. 3.42). In the aluminium oxide intermediate layer
Coated cemented carbide, tc = 20 minUncoated cemented carbide, tc = 40s
vc = 120 m/min, f = 0.32 mm, ap = 2 mm vc = 200 m/min, f = 0.2 mm, ap = 2 mm
Material: X6CrNiMoTi17–12–2
Fig. 3.41 Formation of crater wear on an uncoated and a coated cemented carbide, acc. to
GERSCHWILER [Gers04]
3.7 Wear 83
W
Co
Fe
Cr
Ni
X6CrNiMoTi
17-12-2
WC-7,5Co-
6,8MK TiN
Ti
(C,N) Al2O3 ZrN
X6CrNiMoTi17-12-2WC-7,5Co- 6,8MK Ti(C,N) TiNTiN
Cr
Ni
W
Co
Fe
Fig. 3.42 Verification of the diffusion phenomena between coated cemented carbide and steel
dependant on the coating system on the basis of annealing tests
the traceable amount of these elements is reduced to 0 mass-%. There is above all
no indication that iron, chrome, tungsten or cobalt have diffused through the Al2O3
intermediate layer on the opposite side of that layer.
The observation that, excluding air or water, electrically isolated oxide layers
function as diffusion barriers on cemented carbides as opposed to semiconducting
nitride or carbon nitride layers can be explained as follows. In electrically isolating
oxides like Al2O3 or HfO2, only ion diffusion is possible, no migration of elec-
trons or holes. The result of this is that “component diffusion” becomes impossible.
“Component” refers to an atom or iron, which theoretically can be split into an ion
and an electron. For reasons of electron neutrality, the iron atom cannot “release”
its electrons outside the layer and then migrate through the layer as an ion. The
consequence is that these oxide layers act as barriers as long as there is no way for
electrons to get around it, and the transport of oxygen ions is also slow enough (in
oxides this transport is empirically slower than cation transport by several orders of
magnitude.) In semiconductive or conductive layers however, electrons can migrate
through the layer together with the ions, leading finally to the iron element migrat-
ing through the layer to alloy with the substrate. Conversely, conveyance of Co or W
into the chip is also possible. Diffusive transport is already possible at the tempera-
tures prevalent on the tool during machining of > 800◦C. The diffusion coefficients
are indeed still small, but the necessary diffusion paths are very short.
While diffusion of iron through the cemented carbide layer into the cemented
carbide substrate is indeed possible in the case of cemented carbides coated with
TiN and TiCN, this diffusion is not as significant as is the case for uncoated sub-
strates. As opposed to uncoated cemented carbides, the coat provides the substrate
84 3 Fundamentals of Cutting
with additional protection from the abrasive action of the flowing chip. Especially
because they are much harder than the substrate, cemented carbide layers help
to reduce abrasive wear to a significant extent. Cemented carbide layers thereby
ensure the high performance level of coated tools and allow for much higher cutting
speeds when cutting steel materials in comparison with uncoated cemented carbides
(Fig. 3.41). But as soon as the hard material layer in the contact zone of the rake face
has been rubbed off and the substrate exposed, diffusion processes now once again
taking place between the substrate and the chip contribute, together with the abrasive
effect of the chip on the substrate, to an acceleration of crater wear.
3.7.2.5 Scaling
If we examine a tool after cutting, we can recognize several annealing colours near
the contact zones, indicating scaling (oxidation) of the cutting tool material. Scaling
varies in importance depending on the cutting tool material alloy and cutting temper-
ature (Fig. 3.43). Cemented carbides already begin scaling at 700–800◦C, whereby
Annealing time / h
In
cr
ea
se
in
 m
as
s
/ %
HW-K10
HW-P25
HW-P10
HT-P05
20
15
10
5
0
1 2 3 4 5
HT-P10
HT-P20
Atmosphere: air
Temperature: 1000 °C
Annealing
temperature:
500 °C
Annealing time:
3h exp. to air
HW-K10
HW-K10 HW-P25 HW-P10 HT-P10
Annealing temperature: 1000 °C; Annealing time:3 h exp. to air
Fig. 3.43 Gain of weight during the annealing process in air dependant on the used cutting tool
material
3.7 Wear 85
Flank face
Oxidation
zones
Rake faceOxidation
Cutting tool material: HW-P30
Cutting velocity:
Cutting time: 
Feed:
Depth of cut: 
Material: C53E
vc = 125 m/min
tc = 20 min
f = 0.25 mm
ap = 3 mm
Fig. 3.44 Oxidation zones at a cemented carbide turning tool
those made of pure tungsten carbide and cobalt oxidize more than those alloyed with
titanium carbide or other carbides [Kief65].
Even under the usual cutting conditions, an oxide film is formed on WC tools in
the area of the cutting edge under the effect of the cutting temperatures and atmo-
spheric oxygen. This film covers the areas to which atmospheric oxygen has free
access, i.e. the ends of the contact zones on the flank face, minor flank face and rake
face (Fig. 3.44).
The destructive effect of oxidation on the structure of cemented carbide can
be observed especially clearly on the minor flank face. A complex tungsten-
cobalt-iron oxide is formed that, as a result of its larger molar volume compared
with cemented carbide, develops in a wart-like fashion and can lead to corner
breaking [Köni75].
Scaling is of no practical importance for tool steels and high speed steels since
their heat resistance is exceeded before their surfaces are oxidized more strongly.
3.7.3 Forms and Dimensions of Wear
Wear phenomena that are generated on the cutting part during cutting vary greatly
depending on the type and duration of stress. Figure 3.45 shows wear forms that
occur primarily on turning tools. The cutting part is worn on the rake face (crater
wear) as well as on the major and minor flank faces (flank face wear) (Fig. 3.46).
Depending on the cutting parameters and the combination of workpiece and cutting
tool material, flank face wear tends to be strongest on the edge of the contact zone
on the major and minor cutting edge of the tool, which continues on the rake face.
Such “notch wear” is caused by the abrupt transitions of mechanical and thermal
stress prevalent on the end of the contact area, the abrasive character of the sharp
86 3 Fundamentals of Cutting
Notch wear on minor cutting edge
Cratering on rake face
Plastic deformation of corner of
cutting edge to the point of
breaking 
Layer flaking beneath width of
flank wear land 
Layer flaking due to plastic
deformation of cutting edge
Notch wear on flank face
Wear on flank and minor
flank face
Delamination on cutting edge
beyond contact zone
Smeared material, partly
overlapping the cutting edge in a
palisade-like manner 
Spalling of cutting edge
due to chip impact or
clamped chips
Spalling of cutting edge along
major cutting edge
Notch wear on rake face
Fig. 3.45 Characteristic wear forms at the cutting part during the turning process
Cutting tool material: Ceramic
Material: Nickel-based alloy
Cutting parameters: 
vc = 220 m/min
f = 0.16 mm
tc = 2.3 min
ap = 0.3 mm
Insert geometry:
RCGX090700T00515
Cutting tool material: Cermet
Material: 42CrMo4+QT
Cutting parameters: 
vc = 400 m/min
f = 0.1 mm
tc = 12 min
ap = 0.5 mm
Insert geometry:
SPGN120308
Cutting tool material: PCBN
Material: Nickel-based alloy
Cutting parameters: 
vc = 300 m/min
f = 0.16 mm
tc = 1.6 min
ap = 0.3 mm
Insert geometry:
VBMW160412
Notch wear on:
Minor cutting
edge
Major cutting
edge
Fig. 3.46 Examples for the formation of crater wear, flank face wear and notch wear dependent
on the used cutting tool material and the machined material
3.8 Cutting Theory 87
KB: Crater width
KM: Distance betw. crater
 centers
KT: Crater depth
SVα: Displacement of cutting
 edge toward flank face
SVγ: Displacement of cutting
 edge toward rake face
Wear notches on
minor cutting edge 
Sectional view
A-A
A
crater
Wear notches on
major cutting
edge 
b
Width of flank wear land VB
VB
B 
m
a
x.
C B N
b/4V
B C VB
B
SV
α
A
rε
VB
N
A
KT
KB
KM
Plane PsSVα
SV
γ
Fig. 3.47 Wear forms and measured quantities at the cutting part, acc. to DIN ISO 3685
chip edge or the feed comb and the tool edge as well as direct contact with the
atmosphere (Fig. 3.46).
Figure 3.47 is a schematic representation of the dimensions of wear. In particular,
we distinguish the width of flank wear land VB, the displacement of cutting edge
toward flank face SVα and rake face SVγ , crater depth KT and the crater centre
distance KM from which the crater ratio K = KT/KM is formed.
3.8 Cutting Theory
3.8.1 Shear Plane Theory
The model representation of the shear plane is based on ideal cutting conditions.
Required are:
• ideally sharp cutting edges,
• cutting edges that are not subject to wear,
• the representation of mechanical loads as substitute forces and
• ideal plastic material behaviour.
88 3 Fundamentals of Cutting
Trace of cutting
edge plane
Ps
Trace of tool reference plane 
Pr
Tool
Po ≡ Pf ≡ Pn
Selected point of cutting edge
Workpiece
Trace of shear plane PΦ
y
z
h
γn
hch
Φ
Fig. 3.48 Model representation of the shear plane
The chip formation theory derived from this model (shear plane theory) assumes
that deformation takes place on only one plane, the shear plane. By definition, the
shear plane passes through the selected point of cutting edge just like the tool-in-
hand and tool-in-use reference system (Fig. 3.48).
In the following the shear plane P� will be considered in the tool-in-hand refer-
ence system. This is a plane which is inclined toward the cutting edge plane Ps by
the shear angle F.
The requirements made on process kinematics are fulfilled when:
• tool cutting edge angle κr = 90◦
• tool inclination λs = 0◦
We are thus assuming a free, orthogonal cut. The shear plane model has the
advantage of making complicated cutting processes highly comprehensible.
If the effects of the corner radius and minor cutting edge are negligible in com-
parison with that of the major cutting edge, the theory of the shear plane can also be
used for the bound cut.
With the shear angle we can macroscopically consider material and cutting tool
material properties as well as thermal loads and friction conditions which pre-
dominate during contact of the chip with the rake face. Moreover, the shear angle
establishes the relation between undeformed chip thickness h and chip thickness
hch.
hch = cos (� − γn)
sin�
· h (3.14)
3.8 Cutting Theory 89
By means of the shear angle we can now calculate the surface area of the shear
area A� in the shear plane:
A� = b · h
sin�
(3.15)
The ratio of undeformed to deformed magnitudes is designated in forming tech-
nology as deformation. If the material is compressed and only the change in height
is described, the term degree of compression is also used. This concerns volume
consistency. In cutting theory, the term chip compression λh is used analogously
[Krys39].
λh = hchh =
cos (� − γn)
sin�
(3.16)
The shear angle � is a purely theoretical quantity and should not be confused
with the actual slip direction of the material, even if the shear surface and slip surface
are nearly identical under certain marginal conditions.
In conclusion, to show that the theory of the shear plane makes it possible to
calculate various quantities with complicated relations, three classical applications
of shear plane theory will be described.
3.8.1.1 Calculation of Chip Speed
From simple trigonometric relations (law of sine) in the velocity plan, MERCHANT
calculated the chip speed vch with the help of the shear angle F (Fig. 3.49) [Merc45].
|
vch| = sin�
cos (γn − �) · |
vc| (3.17)
Trace of cutting
edge plane
Ps
Trace of tool refernce plane
Pr
Tool
Po ≡ Pf ≡ Pn
Selectedpoint of cutting edge
Workpiece
Trace of shear plane PΦ
Φ
y
z
h
γn
hch
vch
vc
vΦ
Fig. 3.49 Velocity plan, acc. to MERCHANT [Merc45]
90 3 Fundamentals of Cutting
The continuity equation can also be used to calculate chip speed vch. The idea
behind this is that a mass flow flows through a control surface, in this case the shear
surface. Assuming that the density of the material is not altered by mechanical and
thermal load, we obtain:
m˙ = const. ⇒ |
vc| · A = |
vch| · Ach (3.18)
From which follows:
|
vch| = hhch · |
vc| =
1
λh
· |
vc| (3.19)
3.8.1.2 Calculation of Resultant Force
It is assumed that the shear angle is a function of the shear stress in the shear plane,
which will be shown in the following steps. In the first place we have:
τ� = |
F�|
A�
(3.20)
The component of the resultant force in the shear plane, the shear force F�, can
be represented by the resulting resultant force Fz. With the help of the angle relations
in the circle of THALES we obtain according to Fig. 3.50:
|
F�| = |
Fz| · cos (� + ρ − γn) (3.21)
Trace of cutting
edge plane
Ps
Trace of tool reference plane
Pr
Po ≡ Pf ≡ Pn
Trace of shear plane
PΦ
y
z
γn
Tool
FΦn Fz
FΦ
Fγ
Fγn
Fc
Ff
Φ
ρ − γn
Fig. 3.50 Circle of THALES for the force diagram, acc. to MERCHANT
3.8 Cutting Theory 91
The amount of resultant force can now be calculated from Eqs. (3.20) and (3.21):
|
Fz| = τ�
sin� · cos (� + ρ − γn) · b · h (3.22)
The functional relations of the values of the resultant force components cutting
force and feed force can be taken from the circle of Thales (Fig. 3.50).
|
Fc| = |
Fz| · cos (ρ − γn) (3.23)
|
Ff| = |
Fz| · sin (ρ − γn) (3.24)
If we now insert Eq. (3.22) into Eqs. (3.23) and (3.24), the following is valid for
the values of the resultant force components:
|
Fc| = cos (ρ − γn) · τ�
sin� · cos (� + ρ − γn) · b · h (3.25)
|
Ff| = sin (ρ − γn) · τ�
sin� · cos (� + ρ − γn) · b · h (3.26)
3.8.1.3 Calculation of the Shear Angle
KRYSTOF developed a model for the simple calculation of the shear angle [Krys39].
He made the following assumptions:
• The theory of the shear plane is valid.
• The maximum shear stress leads to material collapse.
In the main axial system, maximum shear stresses occur at an angle of 45◦.
For calculations of the cutting process, the position of the main axial system
is assumed to be approximately in the direction of the resulting contact stress
(Fig. 3.51).
For purposes of simplicity, the principle stress is not represented by the stress
tensor but simply by vector addition. In this case, the friction angle approximates
the orientation of the main axial system.
The function relation between the shear angle, the tool orthogonal rake angle and
the friction angle can be derived from Fig. 3.51. The following is valid for the shear
angle:
� = π
4
− (ρ − γn) (3.27)
MERCHANT has suggested a possibility of calculating the shear angle that makes
the following assumptions [Merc45, Merc45a]:
• The theory of the shear plane is valid.
• The position of the shear plane is determined by the minimum of cutting energy.
The value of cutting force can be read off the circle of THALES in Fig. 3.50:
92 3 Fundamentals of Cutting
Trace of shear plane PΦ
Trace of cutting
edge plane
Ps
Trace of tool reference plane 
Pr
Po ≡ Pf ≡ Pn
Tool
Fz
1
τΦ
Fγ
Fγn
3
4
−
ρ
π
γn
Φ
Fig. 3.51 Principle of the maximum shear stress, acc. to KRYSTOF [Krys39]
|
Fc| = cos (ρ − γn)
cos (� + ρ − γn) ·
b · h
sin�
· τ� (3.28)
The necessary and the sufficient conditions for calculating the position of the
energy minimum are
nec.:
∂Ec
∂�
= 0 suff.: ∂
2Ec
∂�
�= 0 (3.29)
Since the cutting energy is calculated as the product of the cutting force and
the cutting path, and the cutting path is not a function of the shear angle, we can
simplify to:
nec.:
lc · ∂|
Fc|
∂�
= 0 suff.: ∂
2|
Fc|
∂�
�= 0 (3.30)
From which follows for the shear angle:
� = π
4
− 1
2
· (ρ − γn) (3.31)
3.8.2 Application of Plasticity Theory to Cutting
HUCKS and OPITZ developed a shear angle equation by applying MOHR’s slip the-
ory to the orthogonal cutting process [Huck51, Opit53, Mohr06]. As opposed to
the models of KRYSTOF and MERCHANT, this theory makes it possible to take a
material-specific slip angle into consideration instead of an idealized 45◦. It should
3.8 Cutting Theory 93
y
ρ
yFE
zFE
3
γn
Φ3
1
σ y
yMT
zMT
Tool
Po ≡ Pf ≡ Pn
Sli
p d
ire
cti
on
Slip direction
τzy
Trace of tool reference plane Pr
Trace of cutting
edge plane
Ps ≡ Pp
Trace of shear plane PΦ
1,3: Principle normal stress directions
ϕ
ω
τzy
τyz
τyz σ y
Φ1
Fig. 3.52 Surface element at the cutting part, acc. to HUCKS [Huck51]
again be pointed out that the previously known shear angle need not necessarily be
equal to the angle given by the actual position of the slip surface, since the direction
of the shear plane was previously only purely geometrically determined by ideal
considerations. In this case the directional angle of the slip surface towards the tool
cutting edge plane is designated as the shear angle �. First a square surface element
is considered on the rake face that experiences both shear and compressive stresses
from the tool (Fig. 3.52).
The shear stresses also appear on all remaining faces of the quadratic element
as a result of the law of equality of shear stress. Due to the unhindered flow of the
chip over the face there is a very small amount of compressive stress in this direc-
tion, which will be neglected in the following. There are therefore two compressive
stresses and four shear stresses on the surface element under consideration.
If we now determine the principle normal stress with the help of MOHR’s stress
circle, we establish the validity of the following equation:
ω = 1
2
arctan(2 · μ) (3.32)
The angle between the rake face and the principle stress direction is thus deter-
mined by the friction coefficient between the rake face and the chip. This relation
makes it clear, for example, why different cutting speeds yield different shear angles.
The friction coefficient is dependent on the cutting speed and consequently alters the
principle stress direction. The slip direction on the other hand is inclined towards
the principle stress directions by a materially dependent angle ϕ. This angle can be
94 3 Fundamentals of Cutting
calculated from the compressive flow stress σD and the shear flow stress τF with the
help of flow tests and MOHR’s stress circle:
ϕ = 45 − 1
2
arcsin
(
σD − 2τF
σD
)
(3.33)
For an ideal plastic body which is approximately realized by steel at the yielding
point, the flow plane is under 45◦ in relation to the principle stress direction, since
there the shear stress is at a maximum [Huck51].
With this, we obtain the principle stress and slip directions shown in Fig. 3.52
relative to the rake face. With the help of the angles γ , ϕ, and ω, the slip angle �1
can now be calculated:
�1 = ϕ − 12arctan(2μ) + γn (3.34)
Slip angle �3 is in the still undeformed area, which is why only slip angle �1
(which is identical with the shear angle) is significant. The shear angle calculated
here presumes an extrapolation of the stress conditions via the shear plane and the
width of flank wear land as well as the formation of a continuous chip. This approx-
imation was proven to be in accordance with experiments made on materials with
envelope lines that are inclined and parallel to the σ -axis. HUCKS did not experiment
with materials with bent envelope curve [Huck51].
With the help of the shear angle, the resultant forces can now also be calculated.
To do this, the stressesthat are active in the shear plane are multiplied with shear
plane surface, resulting in a normal and tangential force in the shear plane. We
will in this context proceed without the derivation of HUCKS’s formulae. For steel,
which has an envelope line that is parallel to the σ -axis, the following formulae for
cutting and feed force result with width of undeformed chip b and chip thickness h
according to HUCKS:
Fc = τ0 · b · h ·
⎡
⎣ 1
2
√
μ2 + 14
+ cot g(φ)
⎤
⎦ (3.35)
Ff = τ0 · b · h ·
⎡
⎣ cot g(φ)
2
√
μ2 + 14
− 1
⎤
⎦ (3.36)
To sum up, we can say that the highly numerous experimental relations in chip
formation depend only on a few numbers of materials and the friction coefficient.
HUCKS’s work made the fundamentally new discovery that the stress field in the
chip on the cutting edge depends, besides its intensity, only on the friction coefficient
and the normal direction of the envelope line [Huck51].
Chapter 4
Cutting Tool Materials and Tools
Tool change times, and with them both manufacturing times and tool, machine and
labour costs, are affected by wear. Wear is affected in turn by the properties of the
cutting tool materials. Development in the cutting tool material sector is therefore far
from finished, but is constantly aiming both to improve cutting tool materials that are
already established as well as to discover new materials for use in the manufacture
of cutting tools.
Cutting tool materials should have the following properties in order to do justice
to the stresses placed on them:
• hardness and pressure resistance,
• bending strength and toughness,
• edge strength,
• inner bonding strength,
• high temperature strength,
• oxidation resistance,
• small propensity to diffusion and adhesion,
• abrasion resistance,
• reproducible wear behaviour.
Should we consider all these characteristics simultaneously, we are then faced
with the requirement for the “ideal” cutting tool material (Fig. 4.1). However, there
is no one cutting tool material that unifies all the required properties. One rea-
son for this, for example, is the physical opposition of hardness and toughness.
Developments in the cutting tool material sector are concentrating on optimizing
and modifying chemical composition, manufacturing methods, geometry and coat-
ing in order to broaden the areas of application of cutting tool materials and tools in
accordance with the requirements of modern production.
4.1 Overview of Cutting Tool Materials
The cutting tool materials used in cutting can be summarized as (Fig. 4.2):
• tool steels,
• cemented carbides,
95F. Klocke, Manufacturing Processes 1, RWTH edition,
DOI 10.1007/978-3-642-11979-8_4, C© Springer-Verlag Berlin Heidelberg 2011
96 4 Cutting Tool Materials and Tools
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nc
e
1 Considering the hardness
2 Considering the hardness
 and the temperature
DP:
2
Durability and bending-strength
Feed
HSS
Coated HSS
Fine grain
and submicron
grain cemented
carbide
Al2O3 + TiC
Coated
cermet
Coated
cemented-
carbide 
Si3N4 -
ceramic
Al2O3 -
ceramic
Cemented-carbide
on tungsten-carbide-basis
Cermet
Fig. 4.1 Schematic classification of several cutting tool materials
Cold work steels
High speed steels
WC-Co
TiC/TiN - Co,Ni
WC-(Ti,Ta,Nb)C-Co
Oxide-ceramic
Silicon-nitride-ceramic
Mixed-ceramic
Boron-nitride
Diamond
Tool steels
Cemented carbide
Cutting - ceramics
Super-hard cutting
tool materials
Cutting tool materials for chipping
with geometrical defined cutting edge
Fig. 4.2 Classification of cutting tool materials for machining
4.1 Overview of Cutting Tool Materials 97
• cutting ceramics and
• super-hard cutting tool materials made of boron nitride and diamond.
The hardness and wear resistance of the cutting tool materials increases in this
order. On the other hand, their bending strength decreases in this order.
Tool steels include unalloyed and alloyed cold work tool steels as well as high
speed steel. The term cemented carbide refers to conventional cemented carbides
based on tungsten carbide as well as “cermets” based on titanium boron nitride.
Ceramics is the generic term for cutting tool materials made of oxide, mixed and
non-oxide ceramic.
4.1.1 Classification of Hard Cutting Tool Materials
According to DIN ISO 513, the notation and application of hard cutting tool mate-
rials of cemented carbide, ceramic, diamond and boron nitride are classified in
accordance with Fig. 4.3. According to it, uncoated cemented carbides based on
tungsten carbide with a WC grain size of ≥ 1 μm take the abbreviation HW and
with grain sizes of < 1 μm the notation HF. “Cermets” – cemented carbides based on
titanium boron nitride – are notated with HT, coated cemented carbides and coated
cermets with HC. Corresponding abbreviations from Fig. 4.3 are valid for other hard
cutting tool materials based on ceramic, diamond or boron nitride.
The goal of DIN ISO 513, besides that of identifying cutting tool materials, is
above all their assignment to materials, for which they are most suitable for cutting.
Expanding earlier norms, DIN ISO 513 provides six main application groups and
thus six classes of workpiece-materials, which are classified with the code letters
P, M, K, N, S and H as well as by colour (Fig. 4.4). DIN ISO 513 thus maintains
the established code letters P, M and K but now only designate with them the main
application groups that comprise the material groups steel (P), stainless steel (M),
and cast iron (K). The major machining group K included according to the old stan-
dard not only cast iron materials but also a number of other materials. In DIN ISO
513, these now new major application groups were provided with the labels N for
nonferrous metals, S for special alloys and H for hard materials.
Every main application group is subdivided into particular application groups
(Fig. 4.4). These are notated with a code letter for the main application group they
belong to and with an index. The index refers to the toughness and wear resis-
tance of the cutting tool material. The higher it is within each application group,
the lower is the material’s wear resistance and the higher its toughness. The indices
are merely reference numbers indicating a certain sequence. They provide no infor-
mation about the extent of the wear resistance or the toughness of a cutting tool
material. Manufacturers of cutting tool materials should assign their cutting tool
materials to the appropriate application group, depending on wear resistance and
toughness. Examples are HW-P10, HC-K20, and CA-K10. An application group
thus contains comparable cutting tool materials of different manufacturers, although
they may differ in wear behaviour and performance. It is also possible for a cutting
tool material from one manufacturer to be categorized under multiple application
groups if suitable to them.
98 4 Cutting Tool Materials and Tools
Ceramic
Code-letter Group of materials
CA
Oxide-ceramic,
predominantly made from
aluminium-oxide (Al2O3)
CM
Mixed-ceramic, based on
aluminium-oxide (Al2O3),
but also with other
components than oxides
CN
Silicon-nitride-ceramic,
predominantly made from
silicon-nitride (Si3N4)
Code-letter
CC Cutting-ceramic like above,but coated
1) These cemented carbides are also called
 "Cermet"
Cemented carbide
Group of materials
HW
Uncoated cemented
carbide, predominantly made
from tungsten-carbide (WC)
with a grain size = 1 µm 
HT 1)
Uncoated cemented carbide,
predominant
titan-carbide (TiC) or
titan-nitride (TiN) or
both
HC
Cemented carbide like above,
but coated
HF
Uncoated cemented
carbide,predominantly made
from tungsten-carbide (WC)
with a grain size < 1 µm 
CR
Cutting-ceramic,
predominantly made of
Aluminium-oxide (Al2O3),
reinforced
Diamond
Diamond
Code-letter Group of materials
DP Polycrystalline diamond
DM Monocrystalline diamond
Boron-nitride
Boron-nitride
Code-letter Group of materials
BL
Cubiccrystalline boron-nitride
with a low content of
boron-nitride
BH
BC
Cubiccrystalline boron-nitride
with a high content of
boron-nitride
Cubiccrystalline boron-nitride
like above, but coated
Code-letter Diamond category
DP Polycrystalline diamond
DM Monocrystalline diamond
Code-letter Boron-nitride category
BL
Cubiccrystalline boron-nitride
with a low content of
boron-nitride
BH
BC
Cubiccrystalline boron-nitride
with a high content of
boron-nitride
Cubiccrystalline boron-nitride
like above, but coated
Fig. 4.3 Notation of hard cutting tool materials for machining, acc. to DIN ISO 513
4.2 Tool Steels 99
P blue
M
K
N
S
H
yellow
red
green
brown
grey
P01
P10
P20
P30
P40
P50
Steel:
All kinds of steel and
cast-steel,exept stainless
steel witch austenitic
structure 
P05
P15
P25
P35
P45
M01
M10
M20
M30
M40
Stainless steel:
Stainless austenitic and
austenitic-ferritic steel
and cast-steel 
M05
M15
M25
M35
K01
K10
K20
K30
K40
Cast-iron:
Cast-iron with flake–
graphite, cast-iron with
ductile graphite,
annealed cast-iron 
K05
K15
K25
K35
N01
N10
N20
N30
Nonferrous metals:
Aluminium and other
nonferrous metals, non-
metal materials
N05
N15
N25
S01
S10
S20
S30
Specializations and
titanium:
Highly heat resisting
special-alloys based on
iron, nickel and cobalt,
titanium and titanium
based alloys 
S05
S15
S25
H01
H10
H20
H30
Hard materials:
Hardened steel, hardened
cast-iron-materials, cast-
iron for chill-casting
H05
H15
H25
Increasing cutting speed, increasing wear-resistance
of the cutting tool materials
Increasing feed, increasing durability of the cutting tool materials
Main application group Application group
Code-letter Code-color Workpiece-material Hard cutting tool
materials
Fig. 4.4 Application and classification of hard cutting tool materials, acc. to DIN ISO 513
4.2 Tool Steels
DIN EN ISO 4957 defines tool steels as high-grade steels suited to machining and
processing materials as well as handling and measuring workpieces. We distinguish
between cold work steels, hot work steels and high speed steels. Cold and hot work
steels are suited to purposes in which the surface temperature during engagement
100 4 Cutting Tool Materials and Tools
is generally under 200◦C. High speed steels on the other hand can be applied at
temperatures of up to 600◦C. Tools for both cutting and forming are made out of cold
work steels and high speed steels. Hot work steels are used primarily for forming
tools such as forging dies, pressure casting dies or tools for forging machines and
extrusion presses.
4.2.1 Cold Work Steels
Cold work steels obtain their wear and toughness properties from heat treatment,
consisting of heating to austenitizing temperature, quenching in an oil or water bath
(a high cooling speed is necessary, martensite hardness) and tempering (with the
goal of a partial loss of hardness in order to increase toughness).
Cold work steels can be classified as unalloyed and alloyed steels (Fig. 4.5) [DIN
EN ISO 4957, Wegs95]. Unalloyed tool steels (carbon steels) contain up to 1.25% C
Unalloyed cold work steels
Composition Examples
Hammers, hatchets, axes,
scissors, screwdrivers, 
bits
Wood saws, hand saws,
steel blades for segment-
buzz saws
Files, ductors,
chisels, paper knifes
Chisel of all kinds
Files, ductors,
chisels, Paper knifes
Cutting-tools,
shear-blades for
cutting of steel,
broaches, wood-
working-tools
0.80
–
0.90
0.25
–
0.40
0.50
–
0.70
1.20
–
1.35
0.10
–
0.30
0.10
–
0.35
–
0.50
0.40
–
0.40
0.15
–
0.80
0.60
0.42
–
0.47
0.20
–
0.30
0.85
–
1.0
1.7
–
1.9
0.25
–
0.30 0.05
1.10
–
1.25
0.15
–
0.30
0.20
–
0.40
0.5
–
0.8
–
–
0.07
–
0.12
0.15
–
0.45
2.0
–
2.25
0.10
–
0.40
11.0
–
12.0
–
–
–
–
0.6
–
0.8
C45W
1.1730
C85W
1.1830
C125W
1.1563
45CrMoV7
1.2328
115CrV3
1.2210
X210CrW12
1.2436
Alloyed cold work steels
Material grade
Material number C Si Mn Cr Mo V W
Fig. 4.5 Examples for alloyed and unalloyed cold work steels
4.2 Tool Steels 101
and small amounts of Si and Mn. Alloyed tool steels have about 1.25% C as well as
up to 1.5% Cr, 1.2% W, 0.5% Mo and 1.2%V.
The hardness and wear resistance of unalloyed tool steels depend on their marten-
sitic structure. Wear resistance increases with hardness and carbon content; yet
toughness decreases simultaneously and thus the material’s sensitivity during heat
treatment and tool use becomes greater. All unalloyed tool steels are shell hardeners,
i.e. they do not harden all the way through along the entire cross-section but only on
the surface of the workpiece.
The advantages of alloyed tool steels in contrast to unalloyed are the increase in
wear resistance (addition of carbide-forming elements), retention of hardness and
high temperature strength (chrome, tungsten, molybdenum, vanadium alloys) and
in their higher hardness (carbon in solution). Moreover, the critical cooling speed
is lower, allowing for improved hardenability. They can be used at cutting temper-
atures of up to 200◦C. They are used above all in steel machining with low cutting
parameters (reaming, thread die cutting) and to manufacture tools for repair work,
since their cost is lower than high speed steel (HSS) tools because they have fewer
alloy elements.
Due to their low hot hardness, which limits the cutting speeds with which they
can be used, cold work steels are used only rarely for metalworking on machine
tools. Their area of application extends mainly to hand tools such as files, gouges,
reamers or on saw blades for woodworking (Fig. 4.5).
4.2.2 High Speed Steels
A new cutting tool material was introduced for the first time in the year 1900 at the
world’s fair in Paris, with which TAYLOR could realize significantly high cutting
speeds. The productivity of machining processes could thus be increased consid-
erably. This new group of cutting tool materials was called “high speed steels”
(HSS).
High speed steels are high-alloyed steels containing tungsten, molybdenum,
vanadium, cobalt and chrome as their main alloy elements. They have relatively
high bending fracture strength and thus have favourable toughness properties.
In contrast to cold work steels, their matrix is characterized by improved reten-
tion of hardness, and they have higher hardness. Their hardness of about 60–67 HRC
is preserved up to 600◦C in temperature. Due to both this and their machinability,
they continue to have a broad range of application, especially for tools with sharp
cutting edges and small wedge angles such as broaching tools, twist drills, thread-
cutting tools, reamers, milling cutters and turning tools for grooving and parting-off
operations as well as for finishing.
While the hardness of high speed steels are affected by the hardness of the base
material (martensite) and the number and distribution of the carbides, the alloy ele-
ments dissolved in the matrix W, Mo, V and Co, which, partially precipitated as
stable special carbides, are responsible for retention of hardness. Hardness and wear
resistance are increased by the martensite tempered in the matrix and the embedded
102 4 Cutting Tool Materials and Tools
carbides (especially Mo-W double carbides, Cr and V carbides). Carbideformation
and hardening is promoted by additional alloying of chrome.
4.2.2.1 Classification of High Speed Steels
High speed steels are denoted with the letters “HS” and the percents of alloy
elements in the order W-Mo-V-Co (e.g. HS10-4-3-10). High speed steels are clas-
sified according to their W and Mo content into four alloy and performance groups
(Fig. 4.6) [DIN EN 10027a].
Group I includes the steels that contain high amounts of tungsten (18% W).
The type containing Co HS18-1-2-5 has good retention of hardness. Tools made
of this alloy are used for roughing high strength steels, cast irons that are difficult to
machine, nonferrous metals and non-metallic materials.
The material HS12-1-4-5 of the group with 12% W has excellent wear resistance
due to its high V content. Its Co content also gives it high hot hardness and retention
of hardness. This alloy is used in the production of turning tools and profile steels
of all kinds, for finishing tools, high performance milling cutters and automatic
tool machines. They are suitable for machining highly heat-treated Cr-Ni steels and
non-ferrous metals.
For conditioning of steel under...
Composition Symbol
W - Mo - V - Co
< 850 N/mm2 > 850 N/mm2 Roughing Finishing
18% W
HS18-0-1
HS18-1-2-
+
–
–
–
–
+
–
–
12% W
HS12-1- 4 - 5
HS10-4 - 3 - 10
–
–
–
–
(+)
(+)
+
+
6% W 
+ 5% Mo
HS6-5-2
–
–
–
+
–
–
–
(+)
+
–
+
–
2% W 
+ 9% Mo
HS2-9-1
HS2-9-2
HS2-10-1-8
+
–
–
–
+
–
–
–
+
–
–
–
HS6-5-2- 5
HS6-5- 3
I
II
III
IV
5
...average load ...highest load
Fig. 4.6 Alloy and performance groups of high speed steels
4.2 Tool Steels 103
The last two groups are represented primarily by steels containing tungsten and
molybdenum. Molybdenum replaces tungsten from metallurgical points of view and
is more effective at equal masses because it has an approximately twice as large
volume percentage due to its lower density. Steels containing molybdenum are espe-
cially tough. Types containing cobalt are used for machining with simple tools when
robust stresses are expected (drills, turning, milling, planing and broaching tools,
hobs). They are characterized by high toughness, temperature resistance and reten-
tion of hardness. They are universally applicable for roughing and finishing tasks,
especially for highly stressed milling tools.
Steels of both of these groups that have low amounts of or no cobalt are used to
manufacture tools of all kinds. The most important group are high speed steels with
6% tungsten and 5% molybdenum. Usually, the universal or standard type HS6-5-2
is used [Habe88]. As a result of its balanced alloy structure, its high toughness
and good wear resistance, this alloy is very diverse. It is suitable for manufacturing
machining tools for roughing or finishing such as twist drills, milling cutters of all
types, tappers, reamers, broaches etc.
The basic effect of alloy elements in high speed steels is summarized once more
in abbreviated form in the following:
• Tungsten: forms carbides; increases hot hardness, retention of hardness and wear
resistance.
• Molybdenum: forms carbides; improves hardening and toughness, increases hot
hardness, retention of hardness and wear resistance, molybdenum can replace
tungsten (has half its density!).
• Vanadium: exists as the primary carbide VC and increases wear resistance
(finishing).
• Cobalt: shifts the limit of overheating sensitivity to higher temperatures, making
it possible to achieve higher hardness temperatures. Most carbides are dissolved
and the hot hardness increases.
• Chrome: improves hardenability, participates in carbide formation.
• Carbon: the source of hardness in the base material. Increases wear resistance
because it forms carbides.
4.2.2.2 Areas of Application
Figure 4.7 provides a list of the main applications of high speed steels [DIN EN
ISO 4957, Wegs95]. As the amount of alloy elements increases, so does the effec-
tiveness of these cutting tool materials with respect to improved wear resistance
and lifetime. At the same time however, machinability becomes more difficult,
which has negative consequences particularly when manufacturing complicated
moulding tools. Generally, a higher amount of alloy elements means higher tool
costs. The cost effectiveness of a manufacturing process is thus also determined by
the choice of high speed steel.
This use of high-alloyed high speed steels lends itself especially to the solu-
tion of machining problems in which an increase in high temperature strength or
104 4 Cutting Tool Materials and Tools
Steel grade
name according
DIN EN ISO 4957
HS6-5-2
HS6-5-3
HS6-5-2-5
HS10-4-3-10
HS2-9-2
HS2-9-1-8
Material
no.
1.3343
1.3344
1.3243
1.3207
1.3348
1.3247
Main use
Standard high speed steel for all cutting tools for rough turning
or finish turning, taps and twist drills, milling cutters of all 
kinds, pull broaches, reamers, threading dies, 
counterboring drills, planing-tools, buzz saws
Highly loaded taps and reamers, high speed milling cutters, 
broaching-tools, twist drills, 
cutting and shaving wheels
High speed milling cutters, planing- and turning-tools of all 
kinds, highly loaded twist drills and taps, broaches, 
woodworking- and cold working-tools, roughing-tools with
high durability
All-purpose usage for roughing- and finishing-works, 
turning- and highly loaded milling cutters, automated
works, woodworking-tools
Milling cutter, reamers, broaching-tools
End-milling cutter, turning-tools for automated works, 
twist drills, taps
Fig. 4.7 Main applications of the most important high speed steels, acc. to DIN EN ISO 4957
toughness would have a great effect. High speed steels alloyed with cobalt (e.g.
HS6-5-2-5, HS18-1-2-5) are ideal for machining tasks that place increased require-
ments on the high temperature strength of the tools. Steels that contain vanadium in
addition to cobalt, such as the qualities HS12-1-4-5 and HS10-4-3-10, are character-
ized by improved wear properties at increased temperatures stresses and are suited
to machining tasks that place the highest requirements on the wear resistance of the
tools.
4.2.2.3 High Speed Steel Fabrication
High speed steels can be fabricated using cast-metallurgical [Habe88] or powder-
metallurgical methods [Duda86, Beis82]. A comparison of both fabrication methods
is provided in Fig. 4.8. Despite attempts to develop alternative fabrication tech-
niques, the cast-metallurgical fabrication of high speed steels remains the most
economically important.
4.2 Tool Steels 105
Recast
(ESU)
Cast-metallurgical fabrication
Powder-metallurgical fabrication
Hot-
deformation
Housing Hot
deformation
Cold
isostatic
pressing
Sintering
ρ < 100 %
Hot-
deformation
Sintering
ρ = 100 %
Soft-
ening
Water spray 
aerated
powder
GrindingHeat-treatment
Machining
Machining
Machining
Cavity-
pressing
Casting
Gas spray 
aerated
powder
Melting
Hot
isostatic
pressing
Hot 
isostatic
repressing
Fig. 4.8 Production-sequence of the powder- and cast-metallurgical fabrication of high speed
steels
Cast-Metallurgical Fabrication
After melting and casting (1550◦C) the steel in moulds, it is then subjected to block
tempering (900◦C) (homogenization, Fig. 4.9). This is followed by a forging process
(to break up the ledeburite and carbides) with a possible intermediate heating of
the workpiece, as well as by rolling (1200◦C). Softening improves machinability.
The softened condition is in most cases the post practical both for cutting and cold
forming and offers the most favourable output structure for hardening. In order to
avoid cracks or warpage during hardening, itis advisable to anneal the tools with
a minimal amount of stress in order to reduce machining stresses before the last
mechanical processing, especially in the case of tools that are irregular or difficult
to mould.
High speed steels require a highly complex hardening sequence due to the high
amounts of different alloy elements. Conventionally produced high speed steels
thus tend to segregate (demix) during the hardening phase, which is especially pro-
nounced in the case of large cast blocks and high-alloyed steels. Such phenomena
are usually causes of lifetime deviations in later tool use.
Independently of previously applied measures such as improved melting pro-
cess or seeding as well as improved design of the block format, electro slag
remelting (ESR) can further improve the degree of purity and toughness of high
speed steels. In this process, only a small part of the block is fluid during remelt-
ing, not the whole block. In this way, macroscopic demixings such as carbide
agglomerations can be reduced and slag inclusions removed. The disadvantage is
the higher fabrication cost.
106 4 Cutting Tool Materials and Tools
0
200
400
600
800
1000
1200
1400
1600
M
el
tin
g 
an
d 
ca
st
in
g
Bl
oc
k 
tu
rn
in
g
Block
tempering Forging Rolling Softening
Te
m
pe
ra
tu
re
 T
 
/ °
C
Time t /h 
1. 
An
ne
alin
g
2. 
An
ne
alin
g
oil/air
air Air Air AirIf heated
in a vacuum
furnace
Annealing duration
each approx 1–2 h
Time t / h 
Sl
ow
o
ve
n
co
o
l-d
o
w
n Ba
la
n
ci
ng
-te
m
pe
ra
tu
re
Qu
enc
hin
g
Wa
rm
ing
up
Au
ste
nit
isa
tio
n
AnnealingStr
es
s r
elie
f h
ea
t
tre
atm
en
t
Pr
e
pa
ra
to
ry
w
o
rk
Fi
ni
sh
in
g
w
o
rk
s
3. 
An
ne
alin
g
1st preheating stage aprrox. 400 °C
(in salt bath)
D 3rd preheating stage 1050 °C
E Hardening temperature ~1200 °C
F
 Hot bath 500 bis 600 °C
G 50–80 °C
A
B
C
D
E
F
G
Te
m
pe
ra
tu
re
 T
 
 
/ °
C
A 600 to 650 °C
C
 
2nd preheating stage 850 °C 
B
Hardening temp. [°C] 1280 1240 1240 1240 1230
Annealing, each
0.5–1 h
3 ×
560 °C
2 × 570 °C
1 × 550 °C
2 ×
570 °C
2 ×
560 °C
2 ×
540 °C
Code = Grade HS18-1-2-5 HS10-4-3-10 HS12-1-4-5 HS12-1-4 HS6-5-2
Fig. 4.9 Cast-metallurgical fabrication and heat treatment of high speed steels
4.2 Tool Steels 107
The advantage of the ESR process is that the end structure is more even, has
a small amount of segregations and a high degree of purity. The practical conse-
quences of this are reduced warpage in both transverse and longitudinal directions
during hardening and annealing due to the more homogeneous microstructure – an
especially important aspect in the case of tools that are long or are premachined with
a small stock allowance – as well as improved toughness resulting from the higher
degree of purity and the generally more homogeneous structure.
Powder-Metallurgical Fabrication
The starting materials for powder-metallurgical fabrication of high speed steels
is powder obtained by means of gas or water spray aeration of the molten mass
(Fig. 4.8). Gas spray aerated powder is housed and undergoes hot isostatic pressing.
After soft deformation, the semi-finished products are conventionally machined into
tools. Water spray aerated powder undergoes a cold isostatic pressing after soften-
ing, and then the semi-finished products are sintered in a vacuum furnace. Sintering
is not performed until the full theoretical density is reached. Final pressing takes
place later in a forging process downstream. On the other hand, mouldings such
as indexable inserts are hardened in matrix presses. During sintering, the aim is to
reach the full theoretical density. The sintering process can sometimes be followed
by a hot isostatic repressing.
Powder-metallurgically obtained high speed steels have become increasingly
important in recent years. As opposed to cast-metallurgically fabricated high
speed steels, they generally have a somewhat higher alloy content. PM steels are
characterized by a homogeneous structure (no carbide segregation) with an even dis-
tribution of fine carbides. Due to their structural composition, PM steels have better
grindability and higher toughness. With respect to their efficiency as cutting tools,
powder-metallurgically fabricated high speed steels vary to some degree. Numerous
cutting experiments have shown that they are at least equal to conventional high
speed steels with the same nominal composition. PM steels become advantageous
under high mechanical stress resulting from greater feeds, especially when cutting
tool materials that are difficult to machine, such as nickel-based and titanium alloys.
Until now, powder-metallurgically fabricated high speed steels have no mate-
rial indexes of their own. They are designated according to the company that
manufactures them. To simplify their classification, the American nomenclature is
often used, or the standard designation of the corresponding cast-metallurgically
fabricated high speed steel is applied.
4.2.2.4 Heat Treatment of High Speed Steels
HSS tools are subjected to heat treatment in order to give them their required
hardness [Habe88]. First they are hardened (heating and maintaining an austenitisa-
tion temperature followed by high-speed cooling) and then annealed several times.
Figure 4.9 shows a time-temperature chart for the heat treatment of high speed steel
HS6-5-2.
108 4 Cutting Tool Materials and Tools
Due to their poor heat conductivity, high speed steels must be heated very slowly
in order to avoid heating stresses and cracks. Warming up thus generally takes place
in two or three preheating stages at 400, 850 and 1050◦C. After reaching the level
of austenitisation, a temperature of 1150–1250◦C is held for hardening. The goal
of austenitisation is to dissolve as many carbides as possible, whereby the dis-
solution of the secondary carbides increase with a rising hardening temperature
and extended holding times. On the other hand, coarseness must also be avoided.
Depending on the critical quenching speed and on the shape and size of the work-
pieces, quenching can take place in different media, including water, oil, hot baths
(salt or metal molten baths) or gas. Gaseous media include air that is still or in
motion, nitrogen or other gasses. Quenching intensity, which in the case of gasses is
much lower than with liquid media, can be increased by raising the flow speed and
pressure.
After quenching, the microstructure of high speed steels consists of marten-
site, residual austenite and carbides. The hardness of this structure is unsuitable
for cutting tools; due to the high amount of residual austenite and the instability of
the structure, it has insufficient wear resistance, is not dimensionally stable during
use and exhibits an increased fracture tendency. The tools must therefore undergo
annealing after hardening.
High speed steels are annealed at temperatures of 540–580◦C, the temperature
range of secondary hardness. This is defined as the increase in hardness at higher
annealing temperatures after an initial decline in hardness from annealing at low
temperatures. Under normal hardness and annealing conditions, this second increase
in hardness leads to hardness values that are clearly higher than those existing after
quenching. Figure 4.10 is a schematic depiction of the processes that occur dur-
ing the annealing of high speed steels. The cutting tool material hardness resulting
~ 600 
A
S
M
A1
A2
M1
M 2
~ 66 
Annealing temperatureT /°C
H
ar
dn
es
s/
HR
C
Fig. 4.10 Schematic illustration of cutting tool material hardnessresulting from the superposition
of various effects during the annealing of high speed steel
4.2 Tool Steels 109
from the superimposition of various effects is plotted as the cumulative curve S as a
function of the annealing temperature.
As the annealing temperature increases, the hardness of martensite (M) decreases
continuously (M1). The cause of this is precipitation of the carbon trapped in solu-
tion and the associated reduction of crystal lattice tension. The precipitation of the
carbides acts against this process, which leads to the decline in hardness observable
during annealing at up to about 350◦C. Extremely finely distributed carbides pre-
cipitate from the martensite (M) above 350◦C and from the residual austenite (A)
above around 450◦C, causing another increase in hardness (M2/A1) due to precipi-
tation hardening. As a result of carbide formation, the matrix becomes low in carbon
and alloy elements, so that part of the residual austenite is converted to secondary
martensite (A2) during cooling from annealing temperature. All three processes –
carbide precipitation from the martensite, carbide precipitation from the residual
austenite and the conversion of residual austenite into secondary martensite – lead
to the increase in hardness known as the secondary hardness effect. The maximum
of the hardness profile is shifted to higher temperatures (560◦C), and, if austeniti-
sation occurs correctly, hardness can be higher after annealing than in the hardened
state [Habe88].
Until now, heat treatment usually took place in salt baths. Their advantages
are optimal heat transfer and the possibility of executing partial heat and surface
treatments. The important disadvantages are the high costs involved in detoxify-
ing the salts and the necessity of cleansing the workpieces of salt residue after
treatment.
The last few years have seen increasing amounts of heat treatments taking place
in vacuum furnaces. The advantages of this method are above all its environmental
friendliness as well as the clean and shiny workpiece surfaces it yields. Quenching
the workpieces from hardening temperature is achieved by flooding them with nitro-
gen at a pressure of 10 bar. The gas is circulated such that it flows at high speed
through the batch and then through a heat exchanger for recooling. Components are
quenched in this manner to 50◦C.
4.2.2.5 Surface Treatment
The wear properties of HSS tools can be further improved by treating the surface by
nitrating (enriching the surface layer with nitrogen by thermochemical treatment at
temperatures of 500–580◦C), steam annealing (formation of a thin iron oxide layer
on the tool surface in a steam atmosphere at about 500◦C), chrome-plating (deposit
of a hard chrome layer with a thickness of 5–50 μm at temperatures of 50–70◦C)
or coating (PVD coating with thin, highly wear-resistant hard material coats such as
titanium nitride (TiN) or titanium aluminium nitride (Ti,Al)N.
4.2.2.6 HSS Indexable Inserts
Indexable inserts made of high speed steel can be manufactured by precision cast-
ing, by machining pre-products or directly using powder-metallurgical methods.
110 4 Cutting Tool Materials and Tools
Profile milling cutter for rotors
HS10-4-3-10 P20 - P30
Indexable inserts
Material:
rpm:
Feed velocity.:
X5CrNi13–4
n = 65 min–1
v f = 35 mm/min
345
Diameter D/mm
480
HSS Cemented carbide
Cu
tti
ng
sp
e
e
d 
v
c
70
 
m
/m
in 98
 m
/m
in
High
temperature wear
resistance
Sufficiant
ductility
at low vc
Requirements for the
cutting tool material
Fig. 4.11 Fit-for-purpose design of indexable inserts composed of high speed steel and cemented
carbide (Source: Fette)
While pre-products still require further machining, powder-metallurgical fabrica-
tion of HSS indexable inserts for example exploit the cost-intensive alloy elements
much more fully. Besides lower manufacturing costs and more efficient utilization
of material, the most important advantage of sintered HSS indexable inserts is their
increased flexibility in tool use due to the use of indexable insert technology. In
the case of manufacturing processes such as milling or drilling, this technology
allows us to fit tools with various fit-for purpose cutting tool materials (Fig. 4.11).
Indexable inserts made of precision casting have so far gained no importance for
steel machining [Bong91, Wähl86].
4.3 Cemented Carbides
Cemented carbides are composite materials. They are composed of the carbides of
transition metals (fourth to sixth subgroup of the periodic system), which are embed-
ded in a soft metallic binder phase made of cobalt and/or nickel. The carbides are on
the boundary between metals and ceramics. They still have metallic properties (e.g.
electrical conductivity), but they assigned to the non-oxide ceramics as “metallic
hard materials” [Horn06, Salm83]. Additionally, cermets contain titanium nitride.
Hard materials are the bearers of hardness and wear resistance. The binder phase
is responsible for binder the brittle carbides and nitrides to a relatively solid body.
4.3 Cemented Carbides 111
The advantages of cemented carbides include good structural uniformity due to
their powder-metallurgical fabrication, high hardness, pressure resistance and high-
temperature wear resistance. At 1000◦C, cemented carbides have the same hardness
as high speed steel at room temperature. It is also possible to manufacture types of
cemented carbides with various properties by means of intentionally changing the
amounts of hard material and binder [Sche88, Kola92].
4.3.1 Historical Development
In 1927, cemented carbides were first introduced as new high performance cutting
tool materials at the Leipzig Trade Fair under the name WIDIA (Fig. 4.12). This was
a revolutionary development at the time, opening up completely new dimensions in
cutting technology. In contrast to the high speed steel tools used hitherto, cemented
carbides allowed the use of twice or three times the cutting speed with the same
cross-section of undeformed chip. The capacities of the machine tools of the time
were far beneath the potential of the new cutting tools. Materials such as chilled cast
iron, which had been very difficult to cut with HSS tools, could be machined easily
with the new cutting tool material. Initially, WC-Co cemented carbides were used
exclusively to machine cast iron materials. Due to the high amount of crater wear,
these cemented carbides were not suited to machining long-chipping steel materi-
als. This changed with the introduction of titanium carbide as an alloy component.
With cemented carbides containing TiC, similar sensational cutting speed increases
were made possible in steel machining as previously was the case when machining
cast iron with WC-Co cemented carbides. This period also saw the development
of the first cermet. Cemented carbides made for patent reasons without tungsten
based on TiC and Mo2C with Nickel as the binder proved in many cases to be
1923 Patent for the fabrication of WC-Co-HM, K. Schröter
1927 Leipzig Trade Fair: 1st WC-Co-HM, producer Krupp, labeling WIDIA
1928 Carboloy (General Electric)
1931 WC-TiC-Co cemented carbide
1934
1937
1942
Titanit (Plansee), Böhlerit (Böhler)
Coromant (Sandvik)
68/69 Implemention of coated cemented carbide
1970 Fine grain cemented carbide
1973 Spinodal - cemented carbide (1st Cermet with TiN)
since
73/74 Accelerated developing of Cermets
1993 Submicron grain and nano crystalline cemented carbide
TiC-Mo2C-Ni (1st generation of Cermets) producer Plansee, labeling Titanit S
Fig. 4.12 Milestones in the development of cemented carbides
112 4 Cutting Tool Materials and Tools
excessively fragile for the machinery of the time [Sche88, Kola92, Kola93, DRP23,
ÖP31, Häus90, Spri95].
Further development of cemented carbides in the followingyears led to continu-
ous improvement of their composition, production and cutting performance. The
influence of carbide grain size on the properties of cemented carbides was rec-
ognized early on. The relation, whereby cemented carbides can be increased in
hardness only with a reduction in toughness, could be overcome with the devel-
opment of fine grain cemented carbide. By reducing WC crystallite size to under
1 μm, both hardness and bending strength could be increased with the same amount
of binder [Sche88, Kola93, Spri95].
The introduction of coated cemented carbides at the beginning of the 1970s was
another great innovation. The combination of tough cemented carbide substrates
with highly wear-resistant hard material coats led to an enormous increase in pos-
sible cutting speeds and tool standing times. The CVD and PVD methods are the
most important coating process variants today. Multi-layer coats on substrates with
binder-rich rim zones or gradient structures are new developments [Sche88, Kola92,
Gill95].
With the development in 1973 of “spinodal” cemented carbides, the first Cermet,
which contained titanium nitride as a further hard material component, the basic
form of today’s highly efficient cermet was created. Cermets are today among the
high performance cutting tool materials that meet the demands of modern cutting
technology excellently by allowing for the use of high cutting speeds with moderate
feeds and realizing long standing times with a high level of reliability. This is due
to their high chemical stability and high-temperature wear resistance, making these
cutting tool materials especially interesting for cutting operations with high thermal
stress on the cutting edge.
4.3.2 Cemented Carbide Production
Cemented carbides are fabricated using various powder-metallurgical means
(Fig. 4.13). Due to the great variety of shapes that cemented carbide components
can assume, very different forming methods can be used. The most general dis-
tinction is between direct and indirect fabrication, combined fabrication and special
methods (injection moulding). The fabrication method used depends mainly on the
geometry and quantity of the product to be manufactured [Sche88, Kola92].
About two thirds of all cemented carbide products are manufactured by direct
fabrication, primarily indexable inserts. Complicated moulding in small quantities,
such as pistons, screws, rolling rings or matrices are made indirectly, i.e. using
additional machining steps such as cutting, drilling, turning or milling. The start-
ing material is a cemented carbide in a pre-sintered or cold-isostatically pressed
condition, the consistency of which is still chalky [Sche88, Kola92].
The individual components of the cemented carbide are weighed out as powder
to a batch and homogenized in mixers. In the wet grinding following this, the grind-
ing fluid (alcohols, acetone, hexan) should protect the powder from oxidation during
4.3 Cemented Carbides 113
Vacuum drying,
screening
HM-powder
Cold isostatic
compression
Machining
Pressing
additives
Spray drying Granulation
HM-modelling material
Extrusion
HM-injection
moulding material
Injection
moulding
Vacuum drying
dispersing
Vacuum drying
dispersing
Debindering
Sintering
Debindering
Sintering/ 
HIP
Debindering
Sintering
HIP
Grinded
mouldings for
wear protection
Thick beams
Profiles
Drills, milling
cutter
Indexable inserts
Mining tools
Small bearing parts
Short beams
Beams, profiles
d < 22 mm
Drills with
integrated coolant
bore
Small mass
products with
complex geometry
Tungsten-
carbide
WC
Auxiliary
metal
Co, Ni
Alloying elements
TiC, (Ta,Nb)C, Mo2C, VC, Cr3C2
Blending, wet grinding, wet-sieving
HM - Granulate
Semi-finished products
beams, profiles
Finished
products
Coating
Plasticising
Cold isostatic
compression
Powder
compaction
Grinding, separating
Fig. 4.13 Fabrication sequences for the production of workpieces made of cemented carbide
(Source: Widia)
grinding and guarantee optimal dispersal of all the components in the suspension.
After the grinding process is finished, the powder mix is prepared in accordance with
the subsequent shaping method. To produce indexable inserts, the powder is trans-
formed into granulate in order to assure good flow properties and a suitable granulate
size for compression into matrices. The granulate is produced with the help of
spray drying or granulation processes. Shaping the indexable inserts is achieved
by compressing the granulate in matrix presses. The matrix press method allows for
short cycle times and is thus especially suited to producing large quantities [Kola92,
Daub95].
Within the fabrication sequence, sintering process is probably the most important
operation, since it is here that the component obtains the mechanical and technolog-
ical properties that are essential for its functional capability. In principle, sintering
is defined as a thermally activated material transport in which the compressed,
loosely bonded powder material (pressed part, green compact) is increasingly com-
pressed due to diffusion-controlled place-shifting processes (surface diffusion, grain
114 4 Cutting Tool Materials and Tools
boundary diffusion, volume diffusion). It is decisive for fabricating cemented car-
bides that not the entire alloy system is converted to a molten state, but rather (as it
is called in professional jargon) sintered with liquid phase [Sche88, Kola92].
Molten baths first appear at about 1300◦C. Tungsten carbide dissolves increas-
ingly in the molten bath. At a sintering temperature of about 1400◦C, the entire
binder phase, composed of cobalt and tungsten carbide, is molten. The liquid phase
wets the carbide, penetrates into all the pores and causes the hard material particles
to glide together under the influence of surface tension into the smallest space. This
particle rearrangement and re-precipitation process leads to tight packing with mini-
mal surface energy. Depending on the green density and sintering parameters, linear
shrinkage can reach 20%. As the mass cools from sintering temperatures, tung-
sten carbide precipitates again from the liquid binder phase. Up to a temperature of
700◦C, a further thermally activated material transport of tungsten and carbon takes
place in accordance with the solubility of WC in Co. Atoms of the hard material
phase still dissolved after solidification in the binder phase stabilize the cobalt-rich
binder phase in the cubic crystalline crystal lattice structure until room temperature;
otherwise the cobalt converts into hexagonal form below 417◦C [Sche88, Kola92].
4.3.3 Components of Cemented Carbides and Their Properties
• WC: Mono tungsten carbide is the most important hard material phase in techni-
cal sintered cemented carbides. WC is soluble in Co, resulting in the high inner
bonding and edge strength of WC-Co cemented carbides. WC is also even more
wear-resistant than TiC and TaC. The applicable cutting speed is limited at higher
temperatures due to its tendency to dissolution and diffusion.
• TiC: Titanium carbide has a low tendency to diffusion. The consequence of this is
that TiC cemented carbides have considerable high-temperature wear resistance
but little bonding and edge strength. Cemented carbides high in TiC are therefore
brittle and fragile. Their use is preferred when cutting steel materials with high
cutting speeds. TiC can join with WC to make a composite carbide.
• TaC: In small amounts, tantalum carbide has a grain-refining effect, thus improv-
ing toughness and edge strength; the inner bonding strength does not decrease as
sharply as is the case for TiC.
• NbC: NbC has a similar effect as TaC. Both carbides appear as a mixed crystal
(Ta, Nb)C in cemented carbides.
• TiN: Titanium nitride is the property-determiningcomponent in all modern cer-
mets. TiN in steel is even less soluble and is thus more resistant to diffusion
than titanium carbide. Nitrogen causes an increase in wear resistance. In addi-
tion, grain growth is inhibited. Cermets containing nitrogen have as a rule a very
fine-grained structure. In a solid state, TiC and TiN are completely mixable. The
physical properties of cermets are based on those of titanium carbon nitride.
• Co: Cobalt is still unrivalled as a binder metal for cemented carbides based on
tungsten carbide. This is due to the high level of solubility of WC in cobalt and
4.3 Cemented Carbides 115
to the good wettability of tungsten carbide crystals by the molten WC-Co binder
phase [Sche88].
• Ni: due to the improved wettability of hard materials, nickel is used as a binder
for cermets. But since nickel is more easy to deform than cobalt, cobalt is also
added with nickel as a binder in cermets today in order to improve their high
temperature properties.
The properties of the cemented carbide substrate are of key importance with
respect to the wear resistance and efficiency of uncoated and coated cemented car-
bides. Cemented carbides should have high hot hardness and pressure resistance
for the sake of increasing the resistance of the cutting edge to plastic deformation.
They should however also have high bending strength and therefore sufficiently
high toughness and resistance to cracking, crumbling and fracture. In general, tough
cemented carbides are low in hardness and pressure resistance. These opposing cut-
ting tool material properties are influenced decisively by the microstructure of the
base material. The tendency is that fracture toughness increases with the cobalt con-
tent and average grain size, while hardness and pressure resistance are decreased
(Fig. 4.14). As the concentration of composite carbides goes up, fracture toughness
is reduced. Tantalus carbide has a favourable effect on temperature change strength.
Co-Concentration
TiC
TaC
TaC
TiC
WC-Grain size Composite carbide concentration
B
en
di
ng
st
re
ng
th
W
ea
r
re
si
st
an
ce
Cutting tool material attributes
Co-Concentration:
WC-Grain size:
Composite carbide concentration:
Co-Concentration:
WC-Grain size:
Composite carbide concentration:
High
wear resistance
High
ductility
(Ti,Ta)CHa
rd
ne
ss
Pr
es
su
re
re
si
st
an
ce
Fig. 4.14 Influencing factors for wear resistance of cemented carbides
116 4 Cutting Tool Materials and Tools
This is exploited in many cases and, in the case of substrates for milling, the ratio
TiC/TaC is modified in favour of tantalum carbide [Sche88, Schi89, Köni90]. The
high material costs put a tight limit on the use TaC alloys however.
4.3.4 Microstructure
In conventional cemented carbides based on WC, the tungsten carbide is usually in
the form of prisms with triangular bases. As opposed to cemented carbides contain-
ing larger amounts of cobalt, in which crystal growth is less hindered, this crystal
form is less well developed in types that are fine-grained and low in cobalt. Cubic
mixed crystals exist both in both cubical (with rounded edges) and nearly spherical
shape (Fig. 4.15). The carbide skeleton is filled with the binder phase.
The structure of cermet contains only rounded composite carbon nitrides. The
core/shell structure of the hard materials is characteristic of their microstructure. The
causes of this are demixing phenomena in the hard materials (spinodal demixing),
mainly however the dissolution and re-precipitation processes taking place during
liquid phase sintering. By selective solution, the Ni-Co molten bath is enriched
with carbide components. Composite carbides precipitate as a rim zone around
the remaining hard material grains when cooling down from sintering temperature
[Mosk66, Rudy73, Kola89, Kief71, Leop87, Kola89a, Kola93a].
Schematic
Hard material
particles:
WC
(Ti, Ta, Nb, W) C
Binder phase:
Co-W-C-
Mixed crystall
Hard material particles:
TiN
(Ti, Ta, W) (C, N)
(Ti, Ta, W, Mo) (C, N)
(Ti, Ta, W, Mo) (C, N)
Binder phase:
Co, Ni (Ti, W, Mo)
Ni3 (Ti, Al) 
Main component
Schematic
Cemented carbide HW-P10
Chart of reflected electrons
Cermet HT-P15
Chart of reflected electrons
Fig. 4.15 Microstructure of conventional cemented carbides and cermets
4.3 Cemented Carbides 117
The microstructure of the hard material particles is determined by the compo-
nents contained in the original powder. The hard materials can exhibit widely varied
textural structures depending on composition and grain size. If the initial powder
consists of unalloyed binary hard material components, usually titanium nitride or
titanium carbide are in the cores (dark cores). When prealloyed powders are used
consisting of composite carbides or quaternary carbon nitrides, higher concentra-
tions of molybdenum and/or tungsten are also found in the cores (light cores). The
back scatter electron image in Fig. 4.15 shows this varying core/shell structure
very clearly. The initial powder consists in this case of single carbides, compos-
ite carbides and carbon nitrides. The hard material particles with the dark cores
have Ti(C,N) in the centre, which is surrounded by a relatively titanium-rich com-
posite carbon nitride (Ti,Ta,W)(C,N). The lighter (Ti,Ta,W,Mo)(C,N) shell consists
primarily of a titanium-tungsten-molybdenum composite carbide containing a rela-
tively large amount of tungsten and/or molybdenum. The hard material particles in
the light core contain this composite carbide inside, surrounded by a titanium and
nitrogen-rich shell.
4.3.5 Classification of Cemented Carbides
Cemented carbide cutting tool materials can be subdivided into three groups. These
are cemented carbides based on:
• WC-Co,
• WC-(Ti,Ta,Nb)C-Co and
• TiC/TiN-Co,Ni.
The term “cermet” is now used for cemented carbides based on TiC/TiN-Co,Ni.
4.3.5.1 WC-Co Cemented Carbides
The cemented carbides of this group consist almost exclusively of hexagonal tung-
sten monocarbide and the binder phase cobalt. They can contain up to 0.8 mass %
VC and/or Cr3C2 and/or up to 2 mass % (Ta,Nb)C as doping additives to control
structural fineness and consistency (Fig. 4.16).
WC-Co cemented carbides are characterized by high abrasion resistance. Due to
the strong diffusion tendency of tungsten carbide, they are not suitable for machin-
ing soft steel materials. They are mostly used for short-chipping materials, cast iron
materials, non-ferrous metals and non-metals, high temperature materials as well as
in stone and wood processing (Fig. 4.4).
According to previous terminology, WC-Co cemented carbides were subdivided
in accordance with the average WC grain size in the sintered microstructure into
fine (0.8–1.3 μm), submicron (0.5–0.8 μm) and ultrafine (0.2–0.5 μm). Deviating
from this, the standard DIN ISO 513 distinguishes only between cemented carbides
with grain sizes ≥ 1 μm (HW) and those with grain sizes of < 1 μm (HF).
118 4 Cutting Tool Materials and Tools
Cemented carbide
Application group (DIN ISO 513)
Grade WC-4Co WC-6Co WC-9Co WC-12Co
Density
(g/cm3) (ISO 3369)
Hardness HV 30 (ISO 3878)
Pressure resistance
(cylinder test)
(N/mm2) (ISO 4506)
5700 5400 5000 4500
Bending strength
(N/mm2) (ISO 3327)
Young’s modulus
(103⋅N/mm2) (ISO 3312)
Fracture toughness
(N⋅m1/2/mm2)
Poisson-constant
Heat conductivity
(W⋅m–1⋅K–1)
Coefficient of thermal expansion
(293–1073 K)
(10–6⋅K–1)
HW-K05 HW-K10 HW-K25 HW-K40
15.1 14.9 14.6 14.2
1730 1580 1420 1290
1600 2000 2350 2450
650 630 590 580
6.9 9.6 12.3 12.7
 0.21 0.22 0.22 0.22
80 80 70 65
5.0 5.5 5.6 5.9
Fig. 4.16 Composition and attributes of conventional WC-Co cemented carbides, acc. to Kolaska
[Kola92]
Conventional Fine-Grain Cemented Carbides (HW)
Conventionaluncoated WC-Co fine-grain cemented carbides (fine cemented car-
bides) with an average grain diameter of 0.8–1.3 μm (acc. to DIN ISO 513 grain
size ≥ 1 μm) are still widely used in cutting technology applications where high
demands are placed on cutting edge sharpness and toughness, e.g. in steel milling,
finishing, grooving and parting-off or in the manufacture of threads [Kola92].
Submicron and Ultrafine Cemented Carbides
Submicron and ultrafine cemented carbides have until now been the terms used in
literature and practice for WC-Co cemented carbides with an average WC grain
diameter of 0.5–0.8 and 0.2–0.5 μm in the sintered microstructure. The small grain
diameter gives there cemented carbides a special combination of attributes: reduc-
ing the WC crystallite size below 1 μm with equal binder content leads both to
increased hardness and bending strength (Fig. 4.17). This property makes submi-
cron and ultrafine cemented carbides applicable to a wide variety of tasks [Kola92,
Daub95, Köni90a, Köni93, Drey01, Gill01].
Valuable submicron and ultrafine cemented carbides are superior to conventional
fine-grain cemented carbides in hardness, edge strength and toughness. Moreover,
they have a small inclination to adhesion and to wear by diffusion. These attributes
4.3 Cemented Carbides 119
14
12
10
8
6
4
2
0
Turning of white iron
(80 Shore) vc = 16 m/min
f
 
=
 
0.1 mm
ap = 1.0 mm
Indexable insert:
SPGN 120308
Sharp-edged
κ r = 75°
To
ol
 li
fe
 T
 
/ m
in
Standard Fine
grain
Submicron
grain
Standard Fine grain
WC-6Co-Cemented carbide
Submicron grain
(B
): B
en
din
g s
tre
ng
th
N
/m
m
2
15
80
17
80
20
00
(H
)
43
00
(B
)
30
00
20
00
(H
): H
ard
ne
ss
H
V3
0
Fig. 4.17 Microstructure and attributes of fine and submicron grain cemented carbides in
comparison to standard K-class cemented carbides (Source: Widia)
are necessary when hardened materials must be machined with finishing quality and
the smallest machining allowances.
Submicron and ultrafine cemented carbides are used whenever high toughness,
high wear resistance and the highest edge strength is required of the cutting edge,
e.g. for broaching, milling and shaping of heat treated and hardened steels, in cast
iron machining and for machining fibre-reinforced plastics and non-ferrous metals
[Kola92, Daub95, Köni90a, Köni93, Drey01, Gill01].
4.3.5.2 WC-(Ti,Ta,Nb)C-Co Cemented Carbides
The cemented carbides of this group contain, besides tungsten carbide, mixed
carbides (MC) made of titanium, tantalum, niobium and/or zirconium carbide
(Fig. 4.18). Compared with WC-Co cemented carbides, they are characterized by
improved high-temperature properties. This is especially the case for hot hard-
ness and high temperature strength, oxidation resistance and diffusion resistance to
iron materials. Their main area of application is in machining long-chipping steels
(Fig. 4.4). Due to the mixed carbide content, the cemented carbides of this group
can be classified according to their use into two subgroups.
Group A: Mixed carbide content >10–11 mass %. Due to their mixed carbide
content, the cemented carbides of this group are characterized by a high level of
120 4 Cutting Tool Materials and Tools
Cemented carbide
Application group (DIN ISO 513)
HW
- - - - --
P10
HW
P15
HW
P25
HW
P30
HW
M10
HW
M15
Composition (mass percentage)
WC
(Ti, Ta, Nb)C
Co
31.0
60.0
9.0
25.5
64.5
10.0
17.3
72.7
10.0
10.0
78.5
11.5
9.5
84.5
6.0
11.0
82.5
6.5
Density
(g/cm3) (ISO 3369)
Hardness HV 30 (ISO 3878) 1560 1500 1490 1380 1700 1550
Pressure resistance
(cylinder test)
(N/mm2) (ISO 4506)
4500 5200 4600 4450 5950 5500
Bending strength
(N/mm2) (ISO 3327) 1700 2000 2200 2250 1750 1900
Young’s modulus
(103⋅N/mm2) (ISO 3312) 520 500 550 560 580 570
Fracture toughness
(N⋅m1/2/mm2) 8.1 9.5 10.0 10.9 9.0 10.5
0.220.220.230.220.230.22Poisson constant
Heat conductivity
(W⋅m–1⋅K–1) 25 20 45 60 83 90
Coefficient of thermal expansion
(293 K–1073 K)
(10–6⋅K–1)
7.2 7.9 6.7 6.4 6.0 6.0
10.6 11.7 12.6 13.0 13.1 13.3
Fig. 4.18 Composition and attributes of WC-(Ti, Ta, Nb)C-Co cemented carbides, acc. to
KOLASKA [Kola92]
high-temperature wear resistance, a low tendency to diffusion with iron materials
and low abrasion. They are utilized above all in steel and cast steel cutting, with the
exception of rust and acid-resistant steels with austenitic microstructures.
Group B: Mixed carbide content < 10–11 mass %. The cemented carbides
assigned to this group have relatively good high-temperature wear resistance
and abrasion resistance. They are especially suitable for cutting rust, acid and
heat-resistant steels with austenitic structures as well as for alloyed or hard
austenitic/ferritic cast iron materials.
4.3.5.3 TiC/TiN-Co,Ni Cemented Carbides (Cermets)
Cemented carbides based on titanium carbide and titanium nitride with a Ni,Co
binder phase are designated as cermets (formed from ceramic + metal). Present-day
cermets are complex multi-material systems that can contain a number of additional
elements such as tungsten, tantalum, niobium, molybdenum or complex carbides,
from which intermetallic phases are formed during sintering.
4.3 Cemented Carbides 121
Cermet
Application group (DIN ISO 513) HT - P05 HT - P20HT - P10
Composition (mass percentage)
Carbon nitrides
Additive nitrides
Co/Ni
89.0
0.6
10.4
85.7
0.8
13.5
82.3
1.0
16.7
(ISO 3369) 6.1 7.0 7.0
(ISO 3878) 1650 1600 1450
(ISO 4506)
5000 4700 4600
(ISO 3327)
2000 2300 2500
(ISO 3312) 460 450 440
7.2 7.9 10.0
0.21 0.22 0.21
9.8 11.0 15.7
9.5 9.4 9.1
Density
(g/cm−3)
Hardness HV 30
Pressure resistance
(cylinder test)
(N/mm2)
(N/mm2)
Bending strength
Young’s modulus
(103⋅N/mm2)
Fracture toughness
(N⋅m1/2/mm2)
Poisson constant
Heat conductivity
(W⋅m−1⋅K−1)
Coefficient of thermal expansion
(293 K–1073 K)
(10−6⋅K−1)
Fig. 4.19 Composition and attributes of cermets, acc. to KOLASKA [Kola92]
The composition and some characteristic attributes of three typical cermet types
is given in Fig. 4.19 [Kola92]. Compared to conventional cemented carbides,
cermets are characterized by their lower density. Some essential differences to WC-
based cemented carbides are their significantly smaller thermal conductivity coupled
with their greater thermal expansion. Due to their lower heat conductivity, a larger
portion of the cutting heat is dissipated with the chip, resulting in a lower total
heating of the cutting insert. Near the contact zones however, the temperature gra-
dient increases inside the cutting tool material. In association with the large thermal
expansion coefficient, this leads to high tensile and pressure stresses in the cut-
ting tool material. The result of this is that cermets react much more sensitively to
temperature change than conventional cemented carbides. They thus have a strong
tendency to form comb cracks, especially in interrupted-cut machining [Köni90a,
Köni93, Kloc96].
Cermets are very hard, have a low tendency to diffusion and adhesion as well as
a high level of high-temperature wear resistance. Due to their high edge strength,
high resistance against abrasive wear and low adhesion, cermets are especially good
for planing steels. Until now, cermets have been predominantly utilized to machine
122 4 Cutting Tool Materials and Tools
Cu
tti
ng
 s
pe
ed
 v
c
 
/(m
/m
in)
0 0.1 0.2 0.3 0.4
Feed f /mm
0
100
200
300
400
500
Fine finishing
ap < 0.5 mm
Rough finishing
ap < 3 mm
Finishing
ap < 2 mm
Cutting time tc /min
W
id
th
 o
f f
la
nk
 w
ea
r l
an
d 
VB
/m
m
0.05
0.100.20
0.30
0.50
0.40
3 5 10 3020
Cermet
Cemented carbide
HW−P10
Material:
Cutting conditions:
Tool life quantity: 240
f = 0.3 mm Ce
rm
et
Co
at
ed
ce
m
en
te
d 
ca
rb
id
e
vc = 400 m/min
f = 0–25 mm
ap = 1–1.5 mm
ap = 1–1.5 mm
Tool life quantity: 150
C45E
C45+N, Rm = 700 N/mm2 
vc = 400 m/min, f = 0.1 mm, ap = 1 mm
 
vc = 220 m/min
Fig. 4.20 Field of application of cermets during turning steel (Source: Kennametal Widia,
CeramTec)
steels with high cutting speeds and small cross-sections of undeformed chip. The
development of tougher cermet types has led to the expansion of their field of
application to include more average roughing conditions (Fig. 4.20).
Basic areas of application include both turning and milling. Cermets also are
suitable for grooving and thread turning. The high wear resistance of the cutting
edges, in association with the low diffusion tendency and high oxidation resistance,
generally leads to better surface qualities than coated cemented carbides are capable
of producing in planing and finishing. At the same time, due to these properties,
they allow for higher cutting speeds in comparison with conventional cemented
carbides. But they are also suitable for cutting speeds below 100 m/min. Tougher
cermet types, corresponding to the range P15 to P25 of conventional cemented
carbides based on WC-(Ti,Ta,Nb)C-Co, are used successfully in average rough
turning operations and in milling (Fig. 4.21). Because of their clear tendency to
4.3 Cemented Carbides 123
Chamfering of gears
Material:
Tools:
Coolant:
Cutting speed
(m/min)
Tool life
quantity
(pieces) 
Tool costs per
workpiece
*Limited by the nmax
of the maschine
Material:
Tool:
Cutting cond.:
Peripheral down-cut milling
Cermet
Cemented
To
ol
 li
fe
 tr
av
el
 p
at
h 
l f /
 m
0 100 200 300
Cutting speed
vc / (m/min)
100 %
70 %
25 %
200
300
600
HSS -TiN
Cermet
260*
150
Cemented
Carbide -TiN
PVD
50
50CrV4, 250 – 300 HB
End milling cutter,
D = 8 mm, z = 3
fz = 0.025 mm
ap = 8 mm, ae = 1 mm
Dry cutting
16MnCr5
End milling cutter, D = 16 mm, (HSS -TiN)
and D = 18 mm,
(Cemented Carbide, Cermet-inserts)
Emulsion
carbide K25
Fig. 4.21 Milling with cermets (Source: VW-Kassel, Kennametal)
form comb-type fractures, cermets should always be dry-milled [Drey97, Köni90a,
Köni93, Kloc96].
4.3.6 Uncoated Cemented Carbides and Cermets
Uncoated conventional cemented carbides are still used consistently in cutting tech-
nology whenever there are high demands placed in cutting edge sharpness and
toughness properties, e.g. in steel milling, finishing, grooving and parting off opera-
tions or in thread manufacture. They are however in direct competition with coated
cemented carbides and cermets, which are replacing them in an increasing amount
of applications. According to DIN ISO 513, uncoated cemented carbides based on
tungsten carbide take the abbreviation HW and uncoated cermets the prefix HT
(Fig. 4.3).
4.3.7 Function Gradient Cemented Carbides
Function gradient cemented carbides are defined as materials, whose composition
and/or microstructure differ in the rim zone and kernel area. The transition from the
rim zone to the kernel area is continuous [Berg03].
124 4 Cutting Tool Materials and Tools
The goal in the development of function gradient cemented carbides was, and still
is, to improve even further the wear resistance of coated indexable inserts (especially
in the case of dynamic stress) by modifying the rim zone. Such modifications, e.g.
via cobalt enrichening and mixed carbide depletion in a approx. 50 μm thick rim
zone, both cracking and fracture sensitivity of the insert are reduced [Berg97]. Such
measures are, however, only practical when applied to indexable inserts intended
for larger cross-sections of undeformed chip and interrupted cuts. In planing oper-
ations, the strength of cemented carbides is generally completely sufficient. In this
case, a rim zone that takes over or supports the wear-reducing effect of the coating
(mixed carbide enrichening) is advantageous [Berg97]. In the following, three types
of function gradient cemented carbides will be introduced.
4.3.7.1 Cemented Carbides with Rim Zone Free of Mixed Carbides
Gradient cemented carbides are understood as cemented carbides, the rim zone
of which is free of hard and brittle (Ti, Ta, Nb)C mixed crystals (Fig. 4.22)
and consists practically of nothing but tungsten carbide and cobalt up to a depth
of about 50 μm. The cobalt content in the rim zone is higher than within. To
reduce impact sensibility, a coarse WC grain is advantageous. Moreover, nitrides,
which are less hard but more wear-resistant than cubic carbides or carbides of
other transition metals (e.g. zirconium carbides) can also be contained in these
rim zones. In this way, cemented carbide substrates with high hardness but very
tough rim zones of lesser hardness and increased resistance against cracking can be
produced.
The fabrication of rim zones free of mixed carbides can take place in various
ways. The basic principle behind it is the formation of a nitrogen gradient in the rim
zone of the cemented carbides [Berg97].
In the most commonly used method, nitrogen is introduced into the initial
cemented carbide mixture as TiN or Ti(C,N). The titanium nitride decomposes dur-
ing the sintering process. The nitrogen dissolving in the liquid phase diffuses from
the cutting tool material into the surrounding atmosphere. In this way, a nitrogen
concentration gradient arises in the cemented carbide plate from the inside outwards.
The result of this is the formation of a titanium activity gradient. Titanium diffuses
into the interior and accumulates there on existing cubic mixed crystals. The same
is valid for tantalum and niobium. The mixed carbides near the surface gradually
dissolve. On the other hand, cobalt penetrates from inside towards the outside. The
growth rate follows a parabolic temporal law. It increases with a rising nitrogen con-
tent and decreases with a rising titanium content. An edge effect is characteristic of
these processes, the end effect of which is that the zone free of mixed carbides is
thinner on sharp edges than on even surfaces [Berg97].
Other method variants include nitrogen pressure treatment, in which the
cemented carbide is subjected to a nitrogen treatment under increased pressure after
sintering, and the creation of zones free of mixed carbides without nitrogen, in which
the cemented carbide contains a defined excess of carbon and a certain sintering
regime is adhered to during cooling [Berg97, Köni90, Yohe93].
4.3 Cemented Carbides 125
Multilayer
coating on 
gradiented
rim zone
P25
100%
Tool life quantity
360 %
Material = C25
vc = 200 m/min
f = 0.15–0.3 mm
ap = 0.5–1.5 mm
Multilayer coating
gradiented
rim zone
Hardness gradient 10 µm
1000 1200 1400 1600
Hardness HV
Fig. 4.22 Influence of the rim zone on the tool life of CVD tools (Source: Sumitomo)
The thickness of mixed-carbide-free rim zones produced in this fashion is up to
50 μm. The methods described can therefore only be used for indexable inserts that
are already in their final shape after the sintering process. At best a minor cutting
edge rounding can still be undertaken. Therefore, in the case of highly accurate
milling inserts that are ground, such toughness-increasing measures are not used,
although they would be highly desirable. In the case of turning operations with tools
of the application groups P20–P40, indexable inserts with rim zones free of mixed
carbides are widely used [Berg97].
As has been shown in practice, thus designed cemented carbides have signif-
icant advantages in their wear behaviour compared with coated ones. Especiallywhen turning with interrupted cut, tools coated according to this design achieve
significantly longer standing times.
126 4 Cutting Tool Materials and Tools
4.3.7.2 Gradient Cemented Carbides with External Ti(C,N) Enrichment
The hallmark of these cemented carbides is an enrichment of nitrides and carbon
nitrides on the surface. The substrates are first tightly sintered and then impinged
with nitrogen. Diffusion of nitrogen into the rim zone leads to the formation and
enrichment of carbon nitrides high in nitrogen on the surface, the composition of
which becomes more rich in carbon with increasing depth. The thickness of the
surface carbon nitride layer can be varied by changing the treatment time. This
process of enriching the rim zone with Ti(C,N) is called coating the substrate
during the sintering process. The thus “coated” cemented carbides are efficient
alternatives to uncoated and conventionally coated tools. Cutting experiments with
gradient cemented carbides with Ti(C,N)-rich rim zones showed clearly improved
wear behaviour compared with uncoated cemented carbides. They can however also
be conventionally coated. This guarantees that the coating adheres better, espe-
cially when one succeeds in designing the coating and the rim zone such that
neither are clearly delineable from each other after the coating process [Leng06,
Leng04].
4.3.7.3 Gradient Cemented Carbides with Internal Ti(C,N) Enrichment
In this substrate variant, first a rim zone free of mixed carbides is produced. Then
nitrogen pressure is altered at a relatively high temperature so that nitrogen dif-
fuses into the interior through the mixed-carbide-free WC-Co rim zone. Diffusion
of Ti through the WC-Co layer is, however, minimal. The rim zone functions as a
membrane that permits the diffusion within of nitrogen but prevents the diffusion
without of titanium. By means of this one-sided diffusion of nitrogen, a transition
of the WC-Co zone to the base cemented carbide develops, i.e. a zone enriched
with Ti(C,N) a few micrometers beneath the surface in the interior of the substrate.
In milling experiments, inserts with internal Ti(C,N) enrichment exhibited excel-
lent wear behaviour in comparison with TiAlN-coated standard inserts [Leng06,
Leng04].
4.3.8 Coated Cemented Carbides and Cermets
Coating cemented carbides with thin, highly wear-resistant hard material coatings
is one of the milestones in their continued development. Coated cemented carbides
consist of a relative tough base body (e.g. P20, K20) on which is applied a 5–20 μm
thick hard material coating made of carbides (e.g. titanium carbide, TiC), nitrides
(e.g. titanium nitride, TiN), carbon nitrides (titanium carbon nitride, Ti(C,N)) and/or
oxides (e.g. aluminium oxide, Al2O3). Although uncoated cermets are already rela-
tively high in wear resistance, hard material coating also can further improve their
wear performance and efficiency just as with the conventional cemented carbides.
According to DIN ISO 513, coated cemented carbides and cermets have the prefix
HC (e.g. HC-K20).
4.4 Coatings 127
4.4 Coatings
Since about 1968, cemented carbides have been coated with hard materials in order
to improve their wear resistance and thus also their performance capability. Today,
the use of coated cemented carbides and high speed steels for machining various
materials is state-of-the-art.
The high efficiency of coated cemented carbides is made even more clear by
the fact that in the meantime more than 80% of all cemented carbides used in cut-
ting have a highly wear-resistant hard material coating. In turning, the percentage is
about 95% and in milling about 60%.
The primary function of the hard material layer is to inhibit contact between
the material, thereby reducing tool wear, which is caused by adhesion, abrasion,
diffusion and oxidation phenomena (Fig. 4.23). These processes are superimposed
to a large extent and are only partially separable in their causes and in their effects on
wear. But in general, besides abrasion at low cutting speeds, it is above all adhesion
and, at high temperatures, diffusion and oxidation that determine wear on the tool.
With their high hardness and chemical stability, several hard materials possess
the required prerequisites for improving resistance effectively against crater, flank
face and notch wear at both low and high cutting speeds.
While the hard material layer improves wear resistance, lessens adhesion
between the tool and the workpiece and acts as a diffusion barrier, the task of the
substrate is to act as an effective support for the hard material layer and to provide
the composite body (consisting of substrate and coating) with sufficient hot hardness
and, in particular, toughness.
The toughness attributes of a substrate are of decisive importance for applications
with interrupted cut. Due to input and output impacts and associated mechanical and
thermal stress variations, the toughness of the cutting edge is much more responsible
Diffusion
Oxidation
Abrasion
Delamination
Adhesion
Surface effects
Volume effectsStrains
Crack formation
Crumbling
Fractures
Fig. 4.23 Wear phenomena at coated cutting tools
128 4 Cutting Tool Materials and Tools
for its dimensional accuracy than wear resistance. For the interrupted cut, the com-
posite material “coated cemented carbide” must first exhibit sufficient resistance to
cracking, crumbling and fracture (Fig. 4.23). Only when these demands are met
can, analogously to the interrupted cut, wear mechanisms like adhesion, abrasion,
delaminations and – as long as the contact zone temperature and contact time are
sufficient – oxidation and diffusion be effectively hindered.
From this we can derive two requirements for coatings [Köni92]:
• The hard material layer has to drastically reduce the effect of all (if possible)
wear mechanisms involved in wear process.
• The coating process may not reduce the inner bonding strength, i.e. especially the
toughness of the substrate.
4.4.1 Coating Methods
Coating cutting tools can be done chemically and physically. Method variants
include
• chemical vapour deposition (CVD)
and
• physical vapour deposition (PVD).
Figure 4.24 shows an overview of technically established coating methods for the
production of wear-resistant hard material layers. As can be seen, layer deposition
lies within a specific pressure/temperature range for each method and includes some
characteristic coating systems, which are given in bold in the illustration [Berg03].
105
104
105
103
102
101
10–1
10–2
100
400 600 800 1000 1200
Coating temperature T /°C
Pr
es
su
re
 p
 
/P
a
PA-CVD
PVD
MT-CVD CVD
(Ti,Zr)(C,N)
(Ti,Hf)(C,N)
Hf(C,N)
Zr(C,N)
Ti(C,N)
Al2O3 / ZrO2 / TiOx
TiC
TiN
Ti(C,N)
Al2O3
(Ti,Zr)(C,N)
TiN
Ti(C,N)
Al2O3
TiN
Ti(C,N)
(Ti,Al)N
(Ti,Al)(O,N)
MoS2
WC/C
Fig. 4.24 Comparision of pressure and temperature conditions of coating methods with charac-
teristic coating systems (Source: Kennametal Widia)
4.4 Coatings 129
4.4.1.1 CVD Processes
CVD processes are chemical reactions that take place in gaseous phase under rough
vacuum conditions (103–105 Pa) and with the addition of thermal or radiant energy,
forming technically useful solids (hard materials) as well as volatile products. The
chemical reaction is determined by the laws of thermodynamics and thus by the
partial pressure of the gaseous components and temperature.
There are three types of CVD coating procedures:
• HT-CVD (high temperature CVD, 900–1100◦C)
• MT-CVD (medium temperature CVD, 700–900◦C)
• PA-CVD (Plasma activated CVD, 450–650◦C)
A broad palette of hard material coating systems can be synthesized with CVD
methods. The most common coating materials are based on hard materials with pri-
marily metallic bonds, such as TiC, Ti(C,N), TiN, heteropolar(ionic) bonds such
as Al2O3, but also covalent bonds such as diamond [Damm04]. Figure 4.25 pro-
vides an overview of the most common hard materials manufactured with the CVD
process.
Figure 4.25 provides a comparison of some essential physical characteristics
with the properties of tungsten carbide, the main component of classical cemented
carbides.
WC-
6Co TiC TiN ZrC ZrN HfC HfN
Crystal system fcc fcc fcc fcc fcc fcc
Density
(g/cm3) 14.9 4.93 5.44 6.46 7.35 12.30 13.94
Micro hardness
HV0,05
1580
HV30
3100 2300 2800 1700 2600 1850
Young’s modulus
(103⋅N /mm2) 590 451 570 348 495 352 530
Melting point
(°C) 3070 2950 3420 2980 3700 3380
Coefficient of
thermal
expansion
(10–6.K–1)
5.5 7.7 9.4 6.7 7.3 6.6 6.9
Thermal
conductivity
(W⋅cm–1⋅K–1) 80 33 29 38 21 29 22
Colour
G
re
y
M
et
al
lic
gr
e
y
G
ol
d
ye
llo
w
M
et
al
lic
gr
ey
Li
gh
t
ye
llo
w
M
et
al
lic
gr
ey
G
re
e
n
ye
llo
w
α-Al2O3 ZrO2 HfO2
rhombo-
hedral
mono-
clinic
mono-
clinic
3.99 5.7 9.7
2300 1200 1100
380 200
2050 2700
8.0 7.7 5.6
6 (at
1000 °C)
0.6 (at
1000 °C) < 0.1
Un
co
lo
re
d
W
hi
te
G
re
y
w
hi
te
Fig. 4.25 Characteristics of hard materials for coatings in comparison to a cemented carbide of
the application group HW-K10 (Source: Kennametal Widia)
130 4 Cutting Tool Materials and Tools
Because the process is near equilibrium, thermally activated HT and MT-CVD
methods can only be used to produce equilibrium phases. On the other hand, the
low temperature PA-CVD process is also capable of producing metastable phases
such as TiAlN. Among the most technically advanced methods is the deposition of
coating systems with nanostructures. In this way, complexly doped hard material
phases made of Al2O3 or superlattice structures made of TiN/TiB2 can be produced
with a single coating layer thickness of 5 nm [Damm04].
The HT-CVD Process
The high temperature CVD process (HT-CVD) is the classic CVD method for
coating cemented carbides. With it, HM tools can be coated at temperatures of
900–1100◦C. High temperature coatings are characterized by a high level of adhe-
sive strength on the substrate, which is produced by an easily reproducible material
transfer between the surface of the substrate and the reactive periphery [Ever94].
In order to produce a TiC coating, titanium tetrachloride (TiCl4), for example,
is vaporized and transferred with methane (CH4) to a reaction vessel, which can
contain several thousand indexable inserts (Fig. 4.26). Titanium carbide is formed
at temperatures of 900–1100◦C and a pressure below atmospheric pressure in a
hydrogenous environment (H2) in a chemical reaction following this equation:
CH4
H2
H2+TiCl4
Fluid
TiCl4
Evaporator Gas exhaust
Coating furnace
Gas inlet
H2
Fig. 4.26 Schematic illustration of a CVD coating unit (schematic)
4.4 Coatings 131
TiCl4 + CH4 + n H2
1000◦C
−−−−−−−−→
10 − 150 mbar
TiC + 4HCl + n H2 (4.1)
This process allows us to deposit hard materials like TiC, TiN, Ti(CxNy),
(Ti,Hf)(C,N), (Zr/C,N), Al2O3, AlON and others as single coats or in different com-
binations as multilayer coats. The coating material is formed by chemical reaction
from the gaseous phase directly on the surface of the parts to be coated. The reac-
tion products rinse the substrates so that no shading effects arise. Parts with complex
geometries can thereby be coated thoroughly and consistently without difficulty.
Cemented carbides that are coated with the classic high temperature CVD pro-
cess are characterized by high wear resistance due to their relatively thick hard
material coatings (up to 20 μm for turning and up to 6 μm in the case of milling).
The high process temperature is problematic in the case of coating tool and heat-
treated steels. In this case, coating must be followed by rehardening, making the
material vulnerable to unacceptable warpage.
A further disadvantage to the HT-CVD coating process is that the toughness of
the coated cemented carbide body is reduced in comparison to the uncoated sub-
strate. The causes for this loss of toughness in HT-CVD coating are extraordinarily
complicated. They include on the one hand the properties of the substrate to be
coated (chemical composition, crystalline structure, grain size, thermal expansion,
bending and pressure strength, pre-treatment) and process control on the other hand
(gas atmosphere, pressure conditions, temperature/time cycle). Furthermore, pro-
cesses taking place during coating in the rim zone (formation of eta phases), residual
stresses present in the substrate, transition zone (interface) and hard material layer
as well as the microstructure, texture, thickness and adhesion of the applied coating
are all mutually responsible for toughness reduction in HT-CVD coating.
We will now consider more closely processes in the rim zone (eta phase
formation), the “temperature/time cycle” in the coating process and coating thick-
ness based on all the above factors, which offer us many starting points for the
improvement of toughness in coated cemented carbides.
During the high temperature CVD coating process, there is a danger of form-
ing brittle phases in the interface. Cemented carbides of the first generation
exhibited an additional approx. 3–5 μm thick brittle zone with an “eta phase”
(W6Co6C, W3Co3C), caused by decarburization processes in the rim zone [Schi89].
Embrittlement from the eta phase has a negative influence on the toughness of
coated cemented carbides.
Figure 4.27 shows the time/temperature transformation chart for a cobalt alloy
with 5% tungsten and 0.23% carbon. It shows typical coating temperatures and
times for the HT-CVD, MT-CVD, PA-CVD and PVD methods. From this chart,
we can see that in the classic HT-CVD and medium temperature CVD process the
area of eta phase precipitation is passed through/tangent. In contrast, no changes in
constitution are to be expected in the case of low temperature coating methods, as
in the PA-CVD or PVD processes [Köni90].
132 4 Cutting Tool Materials and Tools
Eta'EtaWC
Co3W
200
400
600
800
1000
1200
0.01 0.1 1.0 10 100 1000
Te
m
pe
ra
tu
re
 T
 /°
C
Time t /h
HT-CVD
MT-CVD
PA-CVD
PVD
Co-5W-0.23 C
Fig. 4.27 TTT-chart of a cobalt alloy with typical temperature sequence of several coating
processes (Source: Widia)
The comparative examination of the bending strength of thin, variously coated
WC-Co cemented carbides shows what effect coating temperature and coating
thickness have on these toughness-relevant parameters (Fig. 4.28).
With increasing coating thickness and coating temperature, bending strength is
reduced. The effect is especially negative in cutting operations with interrupted cut,
in which the tool is subject to dynamic stresses. In the case of milling therefore,
cemented carbides with thinner hard material coatings (4–6 μm) than in turning are
used. With these coating thicknesses, toughness loss is relatively low. Reduction of
wear protection due to the reduced coating thickness must be expected.
The logical consequence of these discoveries is that more and more coating
methods have gained significance for coating cemented carbides in the past few
years, which can be used at lower coating temperatures in comparison with the
classic HT-CVD process.
The MT-CVD Process
By using acetonitrile (CH3-CN) instead of a methane (CH4)/ nitrogen (N2) gas mix-
ture, it is possible to lower the coating temperature in comparison with the HT-CVD
method by up to 200◦C [Chat86]. In the 700–900◦C temperature range, carbon
4.4 Coatings 133
Ch
an
ge
s
o
f b
en
di
ng
st
re
n
gt
h
/ %
10
0
10
–10
–20
–30
–40
–50
–60
02 4 6 8
PA-CVD
PVD
550 °C
700 °C 550 °C
700 °C
CVD
CVD
400 °C
Coating thickness / µm
14,5
2,6
F
TiN/Ti(C,N)
TiN
Fig. 4.28 Influence of coating method and coating thickness on the bending strength of coated
cemented carbides (Source: Widia)
nitride coatings of the elements Ti, Zr and Hf can be applied with high deposi-
tion rates. As opposed to the conventional HT-CVD process, the MT-CVD method
has the following advantages:
• Thermal stress on the cutting tool material is lower due to the reduced coating
temperature at simultaneously higher deposition rates.
• The danger of decarburization and thus of embrittling eta phases is lower for
cemented carbide substrates.
• The coatings applied with the MT-CVD method exhibit lower tensile residual
stresses compared with HT-CVD coatings.
Trial results from the longitudinal turning test show the same qualitative depen-
dence on cutting speed for MT-CVD coated cemented carbides as for the uncoated
substrate (Fig. 4.29) [Köni92, VDI3324]. Both MT-CVD coated and PVD coated
cemented carbides are clearly superior to uncoated cemented carbides in per-
formance. In higher cutting speed ranges, the performance of the 3 μm thick
134 4 Cutting Tool Materials and Tools
100000
50000
10000
5000
1000
500
100
60 80 100 200 400300
N
um
be
ro
f i
m
pa
ct
s 
n
Cutting speed vc / (m/min)
Material: 42CrMo4+QT
Rm = 980 N/mm2
Substrate: P30/40
SPUN120312
f = 0.2 mm
ap = 2.5 mm
MT-CVD
Uncoated
Arc-PVD
Fig. 4.29 Influence of coating methods and cutting speed on the toughness behaviour of carbides
in an interrupted longitudinal turning test
Arc-PVD-TiN coating and the 6 μm thick MT-CVD-TiN-Ti(C,N)-TiN coating do
not differ. The PVD coating tends to show advantages at low cutting speeds. It
extends the range of applicable cutting speeds clearly to lower cutting speeds to
a range that is also covered by coated HSS. At low cutting speeds, the cutting edges
succumb to flank face wear and no longer to cracking and crumbling. The causes
for the improved toughness are the unaffected properties of the substrate and the
more favourable compressive residual stresses resulting from PVD coating of the
cemented carbides [Köni90].
MT-CVD coated cemented carbides are used above all in milling. There are
significant performance advantages compared with conventional coatings. The end
of their standing times is no longer brought about by cutting edge crumbling, but
in a reproducible way by a continuous wearing of the flank face. Their excellent
toughness properties make these cemented carbides also ideal for milling with cut-
ting fluids. Fewer comb cracks form and crack development is slower than it is
the case with HT-CVD coatings. Transverse cracks and cutting tool material crum-
bling occur later, resulting in longer tool standing times and higher machining
quality.
4.4 Coatings 135
Residual Stresses in CVD Coating Systems
As a rule, coatings deposited by means of the HT or MT-CVD process exhibit tensile
residual stresses. This is, however, no longer exclusively true. In the meantime, there
are also CVD systems with compressive residual stresses available.
Tensile residual stresses originate when cooling from coating temperature to
room temperature and result primarily from the difference of the thermal expan-
sion coefficients between the substrate and the coating or, in the case of multilayers,
between the individual coatings. In addition, stress-free annealing in the substrate
cemented carbide always occurs at the coating temperatures typical for the CVD
process, leading (as opposed to the low temperature PVD process) to a reduc-
tion of the (compressive) residual stresses introduced, for example, during grinding
[West00].
Coatings with compressive residual stresses, generally familiar in the case of
PVD coatings, have the advantage that crack development is hindered in the hard
material coating, and the fracture strength of the coating/substrate composite is
largely maintained. Compressive residual stresses are therefore the most beneficial
to performance in interrupted turning and in milling in general.
Compressive residual stresses can also be produced when using the CVD method
by a corresponding choice and arrangement of individual coating systems. One
example is the coating sequence TiN-Ti(C,N)-Al2O3. The outside coating composed
of Al2O3 has the smallest thermal expansion coefficient, the inside TiN coating the
largest. After cooling from coating to room temperature, the Al2O3 coating exhibits
a small amount of compressive residual stress. Considerable larger compressive
residual stress is measured with the same coating sequence if it is supplemented
externally with a further Zr(C,N) or Hf(C,N) coating (ratio C:N = 1:1) (Fig. 4.30).
Hf(C,N)
(MT - CVD)
Strain Pressure
10
00
–
10
00
–
20
00
–
30
00
0
17
 µ
m
CVD-Coating system PVD-Coating system
–
10
00
–
20
00
–
30
00
0
Pressure
TiAlN
4 
µm
–
40
00
Cemented
carbide
substrate
Cemented
carbide
substrate
Ti(C,N)
(MT - CVD)
Al2O3
Internal stress (N/mm2) Internal stress (N/mm2)
Fig. 4.30 Influence of coating methods and coating structure on residual stresses in hard coating
systems
136 4 Cutting Tool Materials and Tools
To produce these multilayer systems, the classic high temperature CVD process
is combined with the MT-CVD process, whereby the precipitation of the carbon
nitride coatings takes place with the help of the medium temperature CVD process
[Drey97, West00, Berg03, Berg05].
The Plasma Activated CVD Process
In the PA-CVD process (“plasma assisted” or “plasma activated”) – also known in
the literature as PE, “plasma enhanced” – precipitation of fine-grained hard material
coatings takes place at the, in contrast to HT and MT-CVD methods, much lower
coating temperatures of 450–650◦C. At these temperatures, the thermal energy
alone is insufficient for introducing the chemical reaction needed to form the hard
materials from the gaseous phase. For this reason, additional energy is added to the
process by means of the pulsed plasma of a low pressure glow discharge. In this
way, chemical reactions are made possible that, at thermodynamic equilibrium, are
only possible at much higher temperatures [Tabe89, Köni89].
As in the HT-CVD process, coatings made of titanium nitride, titanium car-
bide, titanium carbon nitride and aluminium oxide can be deposited individually
in crystalline structure by changing the gas composition with individually or in an
arbitrary sequence. Due to the non-equilibrium conditions during deposition, there
is also a possibility of synthesizing metastable systems like (Ti,Al)N using PA-CVD
[Lemm03].
The attributes of the cemented carbide remain largely unaffected by the low-
ering of temperature during coating. The residual stresses are in the compressive
range after coating. PA-CVD coated cemented carbides measure about 30% higher
bending strengths than HT-CVD coated ones.
The properties of PA-CVD coated cemented carbides have a positive effect on
their performance when cutting strong steel materials in interrupted cut. The sensi-
tivity of the composite to comb cracking and failure due to crumbling is much lower
than it is the case for HT-CVD coated cutting tool materials. The application exam-
ple “side milling” proves this (Fig. 4.31). As opposed to the uncoated substrate,
the tool life travel path is slightly diminished after HT-CVD coating due to thermal
influence on the substrate as well as tensile residual stresses in the hard material
coating, while it is over 300% larger after PA-CVD coating.
Modifying the Layer Structure
The layer structure of many multilayer systems usually begins with a thin tita-
nium nitride layer coating the substrate,followed by one or more titanium carbon
nitride intermediate layers. The TiN surface layer functions as a diffusion barrier
during coating and prevents brittle eta phases forming in the interface [Köni90].
The Ti(C,N) layers building upon the TiN surface layer can consist of one layer or
several layers differing in stoichiometry. They act as a kind of “buffer” between the
relatively soft and tough substrate and the much harder and more brittle carbides,
nitrides or oxides of the coating.
4.4 Coatings 137
Cemented carbide P25 / TiN, PA-CVD-coated
Side milling cutter
Material: 60WCrV8,
Cutting tool material: HW-P25
vc = 80 m/min
fz = 0.08 mm
ae = 42 mm
ap = 6 mm
Coating method
To
ol
 
lif
e 
tr
a
v
e
lp
at
h 
L f
 / 
m
PA-CVDUncoated HT-CVD0
4
12
16
8
220 HB
Fig. 4.31 Cutting edge durability of a PA-CVD-coated cemented carbide in side milling (Source:
Widia)
With increasing coating thickness, the wear resistance of the CVD-coated
cemented carbides increases, but their bending strength and toughness is reduced.
The toughness of the coating/substrate composite is affected not only by the thick-
ness of the entire coating, but also to a large extent by the coating material and the
thickness of the individual coating components.
Figure 4.32 illustrates this using the example of a coating made of TiC and
Al2O3. With a constant total coating thickness of 10 μm, the failure rate of the
coated cutting tool material due to fracture clearly increases with the thickness of
the Al2O3 layer [Naka88]. Due to its high chemical stability as a coating material,
aluminium oxide does indeed have excellent resistance against crater wear, but it
only has a small amount of thermal shock resistance and toughness.
In order to reduce the brittleness of coating systems containing aluminium oxide,
they are often designed in the form of fine lamellar, fine grained multilayer coatings
(Fig. 4.32). Such multilayer coatings consist usually of a combination of TiN or
Ti(C,N) with Al2O3 or, as in Fig. 4.32, with aluminium oxynitride (AlON). Coatings
such as these can also be composed of a large number of individual layers, whereby
each layer can be thinner than 0.2 μm. Total coating thicknesses, depending on the
intended use, are around 5 μm in the case of milling and between 12–20 μm for
turning.
Using PA-CVD coating technology, we are now also capable of depositing
nanolayer coatings. These are multilayer coatings with individual layers whose
thickness is given in nanometres, not in micrometers (Fig. 4.33). In the case of
coating systems 1 and 2 shown in Fig. 4.33, sixty 80 nm thick layers and thirty-six
50 nm thick layers are deposited on a top layer composed of titanium nitride. In
138 4 Cutting Tool Materials and Tools
AlON
TiN
AlON
TiN
AlON
TiN
AlON
TiN
Transistion
layers
Substrate
PVD
PA-CVD
CVDBe
nd
in
g
st
re
n
gt
h
Coating thickness
Occurance of failure / (%)
(cutting edge breakage) 
Si
n
gl
e 
la
ye
r
th
ic
kn
e
s
s
Al
2O
3 
/ T
iC
CVD − coating Al2O3 / TiC
Total coating thickness for all tools 10 µm
800 100302010 40 706050 90
6 µm / 4 µm
1 µm / 9 µm
4 µm / 6 µm
3 µm
PVD
CVD
W
ea
rr
e
si
st
an
ce
Coating thickness
Fig. 4.32 Fine lamellar multilayer coating (Source: Widia, Sumitomo)
the case of the coating system TiN-AlN, there are as much as 2000 single layers of
titanium and aluminium nitride of 9 nm thickness that are applied alternately.
Lowering the thickness of the coating increases hardness, toughness and wear
resistance. This effect is based, among other things, on the fact that coating materials
below a certain thickness take on completely new properties. For example, AlN
is, with a hardness of about 1200 HV0.05, relatively soft in comparison to hard
materials like TiC, however it has a hardness of over 3000 HV0.05 when the coating
thickness is below 10 nm.
At high cutting speeds or during dry machining, demands on thermal resistance
of the tool cutting edge are increased. Due to their low thermal conductivity, oxide
coatings are especially suitable to meet these demands. As can be seen in Fig. 4.25,
in addition to aluminium oxide, both zirconium oxide ZrO2 and hafnium oxide HfO2
have very low thermal conductivity. Since both oxides are insufficiently hard by
themselves, they are combined with Al2O3 [West00]. This can be accomplished in
the form of multilayer coating systems with single or multi-phase layers, whereby
the latter – so-called composite coatings – are finely distributed in the Al2O3 base
matrix of a further oxide, e.g. HfO2 oder ZrO2 (Fig. 4.34).
In the case of the Al2O3/ZrO2/TiOx composite coatings shown in Fig. 4.34, the
oxide layer, consisting of three components, is condensed from a gas mixture com-
prising AlCl3, ZrCl4, TiCl4, CO2, H2, N2 and Ar. Al2O3 then exists as a κ-phase and
ZrO2 as a phase mixture in tetragonal/monoclinical (coating temperature > 990◦C)
4.4 Coatings 139
Coating system: TiN-30x(TiCN-TiN)
Coating thickness: 5.2 µm
Single layer thickness: approx. 80 nm
Coating system: TiN-18x(Al2O3-TiN)
Coating thickness: 2 µm
Single layer thickness: approx. 50 nm
Coating system: 1000x(TiN-AlN)
Coating thickness: 9 µm
Single layer thickness: approx. 9 nm
Hardness of coating: 3200HV0,05
Fig. 4.33 Nanolayer created by a PA-CVD coating process (Source: Kennametal Widia)
or monoclinical form (coating temperature < 990◦C) depending on the coating tem-
perature. In the fracture structure image, the ZrO2 particles can be recognized as
finely distributed in the Al2O3 matrix. The highly fine-grained coating structure is
originated by means of a small addition of titanium as doping agent.
Doping layers made of Al2O3 with titanium or boron (or TiN layers with boron)
leads to the formation of very fine-grained coating structures. In the CVD process,
titanium is added to the gas mixture as TiCl4 and boron as BCl3. The doping prod-
ucts TiOx und BxOy, well detectable in the EDX (energy dispersive X-ray) analysis
but still not clearly identifiable as a phase, are effectively dissolved in the Al2O3 lat-
tice. Due to their nanodispersive distribution, they have a positive effect on coating
growth and the grain size of the hard materials forming the coating system [West01,
Kloc01, Berg03]. Doping brings about higher deposition rates, makes it possible to
reduce deposition temperature and clearly improves wear resistance at high thermal
cutting edge loads (Fig. 4.35) [Kath03].
Aftertreatment
Aftertreatment helps to improve further the performance characteristics of coat-
ings, especially in the case of interrupted cut. For example, smoothing the surface
of CVD-coated tool bodies by means of brushing, jet treatment or polishing is
140 4 Cutting Tool Materials and Tools
6 µm
0.5 µm
1 µm
1.5 µm
(12–13 µm)
Ti(C,N) (MT-CVD)
Al2O3 /HfO2 /TiOx
TiN
HfO2
7 µm
5 µm
(12–17 µm)
Al2O3 /ZrO2 /TiOx
Composite
TiCN (MT-CVD)
TiN
9.5 µm
9.5 µm
(16–18 µm)
Al2O3 with dispersed HfO2
TiCN (MT-CVD)
TiN
Fig. 4.34 Fracture structure of CVD-coated cemented carbides with modern coating-design
(Source: Kennametal Widia)
Single layers of TiN and
Ti(C,N) endowed with boron
Ti(C,N)
TiN
Single layers of Al2O3 and TiN
endowed with titan and boron
TiN
MK reduced
rim zone
A
B
A B
Fig. 4.35 Formation of microcrystalline layer structures via doping with titanium and boron
(Source: Ceratizit)
4.4 Coatings 141
the common practice today for removing potential cracking sources. Since brush-
ing only covers the cutting edge area, all-round smoothing with jet treatment is
being given increasingly more preference. It is of particular importance that the jet
treatment, in addition smoothing, alsocauses an alteration of the tensile residual
stresses as a function of the treatment parameters. For example, the Al2O3 layer in
the Al2O3/ZrO2/TiOx coating system has, after jet treatment, compressive residual
stresses (–3.500 N/mm2) in place of tensile residual stresses (1000 N/mm2). Jet
treatment not only improves the residual stress condition, but also reduces the range
within which the wear and standing time of coated hard material indexable inserts
are scattered [Berg03].
Modifying the Substrate Rim Zone
In comparison to uncoated cemented carbides, CVD coatings lengthen standing
times enormously. On the other hand, they also lower the toughness of indexable
inserts, thereby also reducing their capacity to bear high stress peaks during cut
interruptions [Köni90, Köni92]. The reason for this is assumed to be cracks, which
spread from the hard tensile stressed material coating into the cemented carbide
substrate.
Attempts to improve the toughness and resistance of cemented carbides sub-
strates to crack development without reducing their hot hardness led to the
development of substrates with graded rim zone structures [Sche88, Köni90,
Schw89]. The goal was to modify the rim zone of the cemented carbide such that
crack development towards the centre of the indexable insert was hindered or sup-
pressed. One possible solution are cemented carbides, the rim zones of which are
largely lacking hard and brittle (Ti, Ta, Nb)C mixed crystals and consist practically
of only tungsten carbide and cobalt up to a depth of approx. 50 μm (mixed carbide-
free or reduced rimzone, Fig. 4.35). The coating and application of such cemented
carbides is presently state of the art.
Further developments in this area are aimed at establishing the basic properties
of a coated cemented carbide body – hard outside and tough inside – already during
sintering [Leng04]. This would also have positive effects on the following coating
process. The fabrication of function-graded cemented carbides with rim zones lack-
ing mixed carbides and with external or internal Ti(C,N) enrichment is described in
more detail in Sect. 4.3.5.
4.4.1.2 PVD Coating
The early 1980s saw the succession of the CVD process by the PVD process in vari-
ous process variants (vacuum evaporation, Arc-PVD, sputtering). The initial benefit
of these methods was the possibility of coating HSS tools with complex geometries.
In the meantime, these processes are also being used for cemented carbide and cer-
met coating. In addition, decorative layers, oxidation layers and corrosion layers can
also be produced.
142 4 Cutting Tool Materials and Tools
The fundamental differences between PVD coating and the classic HT-CVD pro-
cess can be summarized as follows [Sche88, Schi89, Köni90, Rödh87, Knot89,
Gühr89, VDI3198, VDI3824]:
• The process temperature of 160–600◦C means relatively low temperature stress
on the substrate materials, so that temperature-sensitive substrate materials can
also be coated.
• The bending strength of the substrate remains largely unaffected by the low
coating temperature.
• PVD coatings exhibit residual compressive stresses, which currently limit the
possible layer thicknesses to 3–6 μm. Residual compressive stresses reduce the
risk of fracture formation under shock lead and thermal cycling.
• In the PVD process, workpieces that are to be coated require a very careful
pretreatment of the surface and process control in order to guarantee sufficient
adhesion of the coating. CVD coatings have better adhesion due to the effects of
evaporation and diffusion.
• Because of shadowing effects, consistent layer thicknesses can only be partially
realized by means of very costly devices for rotating the parts to be coated. Inner
contours are as a rule only coatable up to a depth/diameter ratio of 1, since the
layer thickness is diminished with increasing depth.
• High number of potential coating systems and substrate materials.
• Besides stable coating systems such as TiN and Ti(C,N), multi-component
metastable phases can also be deposited under the non-equilibrium conditions of
physical vapour deposition. These phases often exhibit structures that do not exist
under equilibrium conditions. An example of this is the coating system TiAlN.
We can distinguish several different methods of PVD coating:
• methods with thermal evaporation (vacuum evaporation)
• methods with arc evaporation (arc-PVD)
• methods with cathode evaporation (sputtering)
Thermal Evaporation (Vacuum Evaporation)
In vacuum evaporation, usually the coating material is evaporated in a crucible with
a resistance heater or an electron beam gun (EB-PVD) in high vacuum at a pressure
of 10–3–10–6 Pa (Fig. 4.36). The vapour atoms exhibit average free path lengths
of up to several meters at these pressures. Therefore, they generally do not interact
with each other and proceed straight to the substrate. Since the latter is considerably
colder than the vapour, the particles condense on the substrate in accordance with the
principle of water vapour. A columnar fine crystalline layer is formed [Möhw96].
Due to the straight path of the particles, the substrate material must be moved in
the coating chamber in order to avoid shadowing effects and inconsistent coating
thicknesses. Vacuum evaporation allows for high coating rates (coating thickness
increase per unit time) and is used above all for optical applications, in mass pro-
duction for large-scale metallization of films, plastic parts and paper as well as for
coating turbine blades (EB-PVD) [Sche88, Rass96].
4.4 Coatings 143
High vacuum coating plant
 evaporation by:
 - resistance heating
 or by
 - electron beam
Vacuum pump
Energy
supply
Evaporator
Target
Watercooled
recipient
Gas
inlet
Electron beam
Fig. 4.36 PVD-Process – Vacuum evaporation, acc. to Möhwald [Möhw96]
In order to form hard material coatings composed of carbides, nitrides or oxides,
the substrate is reactively vapour-deposited with the help of a reaction gas (N2,
CH2,. . .). [Haef87]. For example, from 2Ti + C2H2 the hard material layer 2TiC and
the gas H2 are formed. In the case of simple reactive evaporation, bonding strength
is relatively low due to the low particle energy of less than 2 eV.
Arc Evaporation (Arc-PVD)
In the arc-PVD method, the target material is vapour-deposited and ionised by
means of an arc in high vacuum. The arc travels either randomly (random arc) or in
a controlled fashion (steered arc) by means of the vaporizing source, usually con-
nected as a cathode (Fig. 4.36). This method is distinguished by its high degree of
ionisation of up to 90% and thus makes it possible to produce high quality coatings
with thick structures and good adhesive properties.
Due to the high energy density in the focal point of the arc, the material is
explosively evaporated and propelled, but molten particles are also torn out at the
same time. The liquid particles that are carried away are found again as drop-
shaped defects (“droplets”) in the layer [Bran05]. These are undesirable because
they adhere to the substrate insufficiently [Möhw96]. The appearance of droplets is
144 4 Cutting Tool Materials and Tools
the main disadvantage of arc evaporation, and since the droplets can lead to pores
in the coating, this method is not suited to depositing corrosion-resistant coatings
[Möhw96]. Droplet emission can be partially reduced by optimizing the process
parameters. Since evaporation by means of cathode arcs requires that the coating
material is electrically conductive, no oxide coatings can be deposited with this
method.
One process variant is the LARC method (Lateral Rotating Arc Cathodes). The
basis of this method are rotating, water-cooled arc cathodes [Damm04]. The mag-
netic field is created by permanent magnets and coils that are shifted and steered
vertically and radially.In contrast to conventional arc technology, fewer droplets
are produced, resulting in smoother coatings [Csel03].
Cathode Evaporation (Sputtering)
In a low-pressure plasma, an inert gas (e.g. argon) is ionized by applying high volt-
age. The positively charged inert gas ions are propelled onto the target switched as
a cathode (coating material) and knock out atoms, atomic groups and molecules of
the coating material via impulse exchange [Rass96] (Fig. 4.37).
Vacuum pump
Additional
magnetic system
Evaporator
(Target)
Plasma-room
Sputtering plant
- without additional
 magnetic field:
 Diodes-Sputter-Source
- with additional
 magnetic field:
 Magnetron-Sputter-Source
Gas
inlet
-1...5 kV
Target
Electrode as
additional
ion source 
+
Fig. 4.37 PVD-Process – Cathode evaporation (sputtering)
4.4 Coatings 145
As opposed to vacuum evaporation, sputtering requires a process gas (usually
argon) as well as a high voltage source. After evacuating the recipient, the process
gas is admitted at up to 0.1–1 Pa pressure. The average free path length of the parti-
cles at this pressure amounts to a few millimetres. At a target voltage of –1 to –5 kV,
a glow discharge plasma is created. The plasma burns between the target which acts
as an anode and the substrate connected to the ground or the recipient wall, thus
making it the anode. The plasma consists of ions, electrons and uncharged atoms
[Möhw96].
This evaporation method is universally applicable, since the material is evapo-
rated not thermally but by means of impulse transfer. Almost all substances [Haef87]
and composites of very diverse melting substances can be evaporated, i.e. not only
metallic, but also isolating layers as well as materials with high melting points can
be deposited.
If not only metals or metal alloys are to be deposited, but also carbides, nitrides,
or oxides, this can be achieved with reactive sputtering. In this case, a reaction gas
(e.g. N2, CH4, O2, H2, . . .) is injected in addition to the process gas. From this
reaction gas are obtained the non-metallic components of the coating (e.g. TiN,
TaC, Al2O3, . . .).
As in the arc method, sputtering also involves the possibility of depositing
metastable coatings (i.e. coatings that do not exist in equilibrium). Because the
process conditions are far from equilibrium conditions, it is possible to freeze
metastable composite phases due to their rapid condensation from the gaseous
phase, thus synthesizing a completely new material type that is often superior to
stable boundary systems [Leye04].
We differentiate between DC sputtering, in which the sputtering material is
electrically conductive, and RF sputtering, in which there is a high-frequency
AC field that is used for non-conductive materials. The use of direct current
(DC) plasma, high-frequency (RF) plasma and the combined use thereof as well
as a nearly unlimited number of possible coating materials making sputtering
enormously flexible with respect to the development of application-specific com-
ponent coatings [Rass96]. Although RF sputtering has important advantages when
atomizing insolating materials, this technology is of lesser importance in the
creation of wear-protection coatings due to the appreciably higher capital costs
in comparison with DC sputtering and the significantly lower deposition rates
[Lemm03]. The disadvantages of RF sputtering can be surmounted by using pulsed
DC sputtering (MF: medium frequency sputtering, up to 50 kHz). The coat-
ings produced with sputtering at low vapour-deposition rates have good adhesive
properties.
Magnetron Sputtering
In comparison to pure diode sputtering, magnetron sputtering involves the arrange-
ment of permanent magnets behind the target (Fig. 4.37), the magnetic field lines
of which vertically penetrate the cathode plate and the target affixed onto it. Due
to the overlapping of magnetic and electric fields, the electrons in the plasma are
146 4 Cutting Tool Materials and Tools
directed on extended spiral-shaped paths. In this way, the probability of ionisation
is increased as well as the deposition rate as the discharge voltage is reduced. It
was the invention of magnetron sputtering that has first made possible an economic
utilization of the sputtering effect for coating deposition and thus the commercial
deposition of wear-protective layers by means of sputtering technology [Lemm03,
Bobz00].
High-Ionisation Pulsing
Rising interest in isolating, mostly oxidic wear-protection coatings and the necessity
of increasing ion density in the area of the substrate in order to deposit layers of
dense, optimized morphology have led to the development of pulsed DC plasmas.
With this technology, not only coating systems usually reserved for RF technology
can be deposited, but completely new coating systems can also be synthesized due
to the extremely enhanced ion currents [Lemm03, Leye04].
Pulse technology makes it possible to deposit conductive and electrically isolat-
ing layers in almost arbitrary stoichiometry as well as nano-structured single and
multi-phased layers with new property combinations.
Figure 4.38 shows a (Ti,Al)N coating created with conventional DC technol-
ogy in comparison to one made with the help of pulse technology. One can clearly
recognize the alteration of the coating micromorphology towards finer crystalline
structure made possible with the increased plasma ionisation of pulse technology
[Lemm03]. Pulse technology thus also creates the possibility of depositing crys-
talline γ-Al2O3 at substrate temperatures of < 500◦C or producing (Ti,Al)N layers
with a AlN content of up to 66 mol% (Fig. 4.39).
The coating process with pulsed plasmas is also known as high ionisa-
tion pulsing (H.I.PTM). We distinguish between three process variants: unipolar,
asymmetrical bipolar and bipolar two-cathode magnetron sputtering [Lemm03,
Erke05].
Fig. 4.38 Fracture pattern of (Ti,Al)N-layers on cemented carbide substrates, deposited with
conventional DC-technique (left) and using bipolar-pulse-technique (right) (Source: CemeCon)
4.4 Coatings 147
900
800
700
600
500
400
300
200
100
TiN AlN
xAlN
T 
/ °
C
(Ti,Al)N+ (Al,Ti)N
(Ti,Al)N
(NaCl structure)
(Al,Ti)N
(wurtzite
structure)
0,2 0,4 0,6 0,8
Fig. 4.39 Metastable phase
diagram of the system
TiN-AlN, acc. to Cremer
[Crem98]
Ion Plating
With respect to process control, ion plating is defined as a variant of PVD coating,
in which the condensing layer is altered by means of an additional electric field
applied to the substrate. This method variant is independent of the type of process
and physical principle involved in creating the metal vapour. The prerequisite is a
sufficiently high number of ionised steam and process gas atoms.
In ion plating, the substrate is coated with a negative voltage or “bias voltage”.
The metal vapour is ionized by means of electrodes arranged in a gas space and
electromagnetic fields. Part of the ionised particles are accelerated towards the sub-
strate (Fig. 4.40). Ion plating thus involves surrounding the substrate with gas ions.
By bombarding the growing layer with ion, coating properties such as density and
residual stress are modified and adhesion to the substrate is improved. The resultant
coatings are very adhesive and tight. All PVD methods in which a bias voltage is
used to modify the coating are termed “ion plating” [Bobz05].
In order to produce carbides, nitrides or oxides for coating purposes, here too
a reactive gas is fed into the chamber. This method is called reactive ion plating
[Möhw96, Haef87]. Ion plating can also be utilized for purifying and activating the
component surface prior to coating (also called etching) or for activation during
coating [Rass96].
A variety of ion plating plant designs have emerged as practicable coatingvari-
ants. For tool coating, arc ion plating (AIP) and ion plating with low-voltage arc
discharge are among the process variants that have become the most significant in
tool coating.
148 4 Cutting Tool Materials and Tools
+ 10...5000 V-
Vacuum pump
Magnets for stabilisation
rather leading of the arc
Vaporising via vacuum arc vaporiser:
Contact ignition device
Rotating arc plasma channel
Fig. 4.40 PVD-Process–Ion plating, acc. to MÖHWALD [Möhw96]
Arc Ion Plating (AIP)
In the AIP process, the target material is vaporized and ionised by means of an arc
in high vacuum. This method is characterized by the high degree of ionization of up
to 90% and can thus be used to produce high quality coatings with dense structures
and good adhesion.
The pre-voltage applied to the substrate accelerates the ionised particles, which
impact the substrate at high speed. Their energy is partially converted into heat,
which can result in high coating temperatures (approx. 550◦C) in the substrates.
Magnetron Sputter Ion Plating (MSIP)
MSIP guarantees a stable plasma and high atomization rates [Bobz00]. Coatings
deposited with MSIP exhibit very good mechanical and optical properties. Nearly all
materials are depositable with this method. Since the ionisation degree of the atom-
ized particles is under 5%, the substrates are heated to a much smaller degree than
in the AIP process, so that plastics and materials with low annealing temperatures
can also be successfully coated [Bobz00].
4.4 Coatings 149
Low-Voltage Arc Coating
Ion plating with low voltage arc discharge is characterized by a low acceleration
voltage, with which an electron jet is created from a low-voltage arc discharge. This
jet is directed towards the target (anode) in a water-cooled crucible. The vapour
made of the coating material arising from the electron bombardment is about 50%
ionised and, since the substrate is applied with a voltage of –200 V, directed towards
the coating target. This method is ideal for depositing nitride or carbide layers by
introducing the corresponding reactive vapours [Haef87].
No droplets are formed when ion plating with low-voltage arc discharge is used
for coating. This method thus makes it possible to deposit very smooth layers
(Rz = 0.1 μm) and is thus suitable for coating mirror-polished surfaces without
the need for post-polishing.
The Use of PVD-Coated Cemented Carbides at Low Cutting Speeds
Broaching soft and hardened steels offers an impressive example for the cutting
edge durability of PVD-coated cemented carbides at low cutting speeds [Köni92a].
Characteristic of the broaching process are the generally low applicable cutting
speeds, usually 1–25 m/min depending on the broaching machine, cutting tool
material and the target material.
Built up edge formation is characteristic of machining relatively soft or tough
steels at low cutting speeds. These lead to dimensional and shape inaccuracies as
well as to reduced workpiece surface quality. One measure for reducing built up
edge formation among others is the use of broaching oils.
By using PVD-TiN-coated cemented carbides in place of uncoated ones, built up
edge formation is almost completely suppressed due to lowered adhesion between
the flowing chip and the cutting tool material in the entire cutting speed range under
investigation of vc = 10–63 m/min (Fig. 4.41). Furthermore, the use of coolants – a
constant ecological problem whose maintenance and disposal are ever costlier – is
superfluous.
4.4.2 Specific Properties of Hard Material Coatings
The most commonly utilized wear-protection coatings in metal machining com-
mercially are TiN, TiC, Ti(C,N), Ti(Al,N), CrN and Al2O3. Besides these coating
systems, those that are characterized by a low friction coefficient such as
DLC (diamond-like carbon) and MoS2-coatings are successfully used for wear-
protection. Due to the unsurpassed hardness of diamond, CVD diamond coatings
are also being increasingly used [Lemm03].
Different coating systems can be deposited as monolayers (e.g. TiN, Ti(C,N),
Ti(Al,N), (Ti,HF,Cr)N) or, to improve their wear-protective effect, combined as
multilayers with single layer thicknesses in the micrometers (e.g. TiN-Ti(C,N)-
Al2O3-TiN, (Ti,Al)N-WC/C).
150 4 Cutting Tool Materials and Tools
vc = 63 m/min broaching oil vc = 63 m/min dry
Material: 17CrNiMo6+FP
Cutting tool material: HW6Co
Chip thickness: h = 0.1 mm
vc = 10 m/min dry vc = 63 m/min dry
Uncoated
PVD-TiN-coated
Tool geometry:
5°
αo γo λs rβ
9° 0° 35 μm
0.2 mm
Fig. 4.41 Coating prevents formation of built up edges
The primary goals of multilayer technology are not only to improve the tough-
ness, hardness and wear resistance of the coating but also to increase the total
thickness of the coating. The transitions that are formed between layers of different
materials place limits on shifting, which leads to a large increase in strength, espe-
cially in the case of coatings with very diverse Young’s modulus. For cracks that are
growing in the coating, transitions between the individual layers signify barriers at
which they are forced into energy-dissipating detours, slowing or even stopping fur-
ther growth (Fig. 4.42). The multilayer structure leads, in comparison to an equally
thick monolayer coating, to lower residual stresses. Multilayer coatings also make it
possible to integrate layers with different properties within one coating system, thus
combining the functionalities of diverse coating materials. Such specific coating
functions include, for example, high thermal stability, high hot hardness or the low
friction resistance of intermediate or cover layers [Krie01, Lugs03]. Larger layer
thicknesses mean a larger coating volume that has to be worn. From this fact results
the generally longer standing times as well.
One further development in multilayer technology is the deposition of nanolay-
ers or coatings with superlattice structures. These are coating systems consisting
of a large number of extraordinarily thin layers, the thickness of which amounts to
4.4 Coatings 151
Coating with low
internal residual
stresses
Nanocrystalline
structures
Crack growth
along the
boundary
surfaces
Crack growth
is hindered
. Balance between high residual compressive stresses (bad coating adhesion)
 and low blank (no obstruction of crack growth) ones 
Multilayer Coatings with
high residual
stresses 
Crack growth
down to the
substrate 
Crack growth
along the grain
boundaries
Substrate Substrate Substrate Substrate
. Ductility of the coating is as important as the hardness of the coating
Fig. 4.42 Schematic diagram of the crack growth in hard coatings depending on structure and
properties (Source: Oerlikon Balzers)
only a few nanometres (Fig. 4.43). In contrast to conventional, usually columnar
layers, coating materials deposited as nanolayers are characterized by their signifi-
cantly higher hardness. As has already been explained in the context of the PA-CVD
coating process, this is based on the fact that the hardness of a material increases sig-
nificantly below a certain layer thickness. This increase in hardness is explained by
the phenomenon that diminishing layer thicknesses have altered crystal lattices and
thus Young’s modulus as well. For example, AlN has a hexagonal lattice structure at
layer thicknesses > 10 nm and a cubic structure when the layer thickness is < 10 nm
[Csel03a].
Another way to improve a coating’s properties is to produce so-called nanocom-
posites. These are noncrystalline isotropic multiphase systems, in which two
mutually insoluble phases (e.g. Al, Ti, Si) are deposited during the coating pro-
cess on the tool surface. Examples of this are the embedment of nanocrystalline
cubic TiN in an AlN matrix, of nanocrystalline TiAlN/AlCrN in an amorphous
Si3N4 matrix[Csel04] or of nanocrystalline (Ti,Al)N into a matrix composed of
(Al,Ti)N. The nanocomposites can be deposited as monolayer or multilayer coating
systems. The boundary layers/grain boundaries in the nanolayer/nanocomposite lay-
ers are energy-dissipating barriers to cracks. Cracking and the speed of crack growth
are thereby reduced (Fig. 4.42). Both nanolayer and nanocomposite layers are thus
characterized not only by extremely high hot hardness and high-temperature wear
resistance but also by favorable toughness attributes.
With respect to the structure of hard material coatings, we differentiate between
monophase layers (e.g. TiN), multiphase layers (e.g. TiN + Ti2N) and graded layers,
i.e. layers the chemical composition of which exhibits a gradient (e.g. (Ti,Al)N
layers with an increasing Al content) [VDI3824].
152 4 Cutting Tool Materials and Tools
Graded layer
Ti(C,N)-Ti(C,N)
(Ti,Al)N-(Al,Ti)N
Monolayer
(TiN, Ti(C,N), (Ti,Al)N, etc.)
Multilayer with functional
intermediate layers
(TiN-Ti(C,N)-Al2O3-TiN,
TiN-Ti(C,N)-Al2O3-Zr(C,N), etc.) 
Nanolayer (Superlattice)
thickness of one layer:
a few atom layers to 100 nm
(Ti,Hf)N-CrN
Hard layer with shear-flexible cover
(TiN + MoS2,(TiAl)N + WC/C, etc.) 
Hard layer + low-friction
amorphous carbon layer
(DLC-layer)
(a-C, a-C:H, a-C:H:Me)
Super hard layer
(CVD-Diamant, Bornitrid)
Nanocomposite
(nc-TiN/a-AlN,
nc-(Ti,Al)N/a-Si3N4) 
Fig. 4.43 Schematic diagram of the structure of current CVD and PVD layer systems
The properties of hard material coatings are determined by their chemical
composition and structure, which depends on the conditions of their deposition.
Figures 4.25 and 4.44 show some characteristic values for layer properties for some
exemplary hard material coating systems produced by PVD and CVD processes. By
changing the deposition conditions, such properties as chemical composition, mor-
phology, structure, texture, residual stresses and thus microhardness, the thermal
expansion coefficient and oxidation resistance can be altered within certain lim-
its. The data presented can thus only be considered as guides for possible areas of
application. Besides the hard material coatings here presented and explained in more
detail below, there is a number of other hard material layers used for tool coating
that are not more fully described here.
4.4.2.1 Titanium Carbide Coatings (TiC)
Coating tools with hard materials began with the deposition of TiC. Due to the high
hardness realizable with titanium carbide (3100–3400 HV0.05), it provides more
effective protection against abrasive wear than TiN. On the other hand, its tendency
to diffusion is a bit higher than that of TiN and Al2O3 due to the relatively small
enthalpy of formation. TiC′s resistance to flank face wear is thus higher than that of
TiN, while its resistance to crater wear is lower. Its oxidation wear is also the low-
est of the layers shown in Fig. 4.44. TiC coatings are chiefly applied using the CVD
process. For the most part, TiC is used as a monolayer when abrasion is predominant
or in multilayer coatings together with TiN, Ti(C,N) or Al2O3 [VDI 3824].
4.4.2.2 Titanium Nitride (TiN)
Since 1980, tools were already being coated with titanium nirtide. TiN is the
most frequently employed hard material for coating cutting tools today. The coat-
ing material is an intercalation compound composed of titanium and nitrogen.
4.4 Coatings 153
TiN TiCN TiC TiAlN
PVD/CVD PVD/CVD CVDProduction process
2300 3000 3100
Coating thickness/µm 1 to 5 1 to 5 1 to 5
Microhardness/HV 0,05 1)
> 450 > 350 > 350
+++ ++ +
Oxidation temperature/°C 2)
++ +++ +++Resistance to abrasion
++ + +
Resistance to wear due to
diffusion
(to steel) 
3000
Resistance to wear due to
adhesion
(against steel)
PVD
1 to 5
> 700
++++
+++
+++
1) For the microhardness, mean values are indicated. They are obtained from reported measured
values resulting from different compositions, coating thicknesses and internal stresses.
2) The oxidation temperature is the temperature at which oxidation of the coating material begins,
considerably affecting the characteristics of the coating.
3) Al2O3 is already an oxide.
4) Being poor conductors of heat, the coatings act as a thermal barrier to the heat produced during
metal-cutting so that most of the heat can be removed through the chip.
5) As the aforementioned hard coatings themselves do not corrode, they protect the basis material
from corrosion. (Leaks in hard coatings may lead to the development of local galvanic elements and
to pitting corrosion).
++ ++ + ++
Thermal barrier effect 4) 
+ + +
Protection of basis material
against corrosion 5) +
CrN Al2O3
PVD
1900
1 to 10
> 600
+
++
++
2100
HV 0.1 
CVD/PVD
1 to 5
–
3)
+++++
++
+++
++ +++
++ +
Fig. 4.44 Quantities Characterising the coating and the coating performance, acc. to VDI 3824/1
Due to the high interaction between the metal and nitrogen atoms, this com-
pound is highly stable. With an enthalpy of formation that is almost twice as
high as that of TiC, TiN is thermodynamically more stable, and thus more diffu-
sion resistant and less inclined to adhesion. Therefore, TiN’s resistance to crater
wear is higher than that of TiC. TiN coatings are characterized by a high level of
toughness.
TiN can be deposited with both the CVD and PVD process. The characteristic
colour of TiN is goldish yellow. Silver-hued TiN coatings consist of a phase com-
posite of Ti2N and TiN, whereby the Ti2N phase is dominant. These coatings are
harder but also more brittle than pure TiN layers. They are most commonly used
when abrasive wear predominates [VDI 3824].
154 4 Cutting Tool Materials and Tools
4.4.2.3 Titanium Carbonnitride Coatings (Ti(C,N))
TiC and TiN can be mixed at any ratio. The properties of titanium carbonnitride are
adjustable by varying the C/N ratio. With increasing carbon content, the colour of
the Ti(C,N) layers change from coppery to violet, bluish-grey to grey [VDI 3824].
Titanium carbonnitrides are used industrially both as hard materials in cemented
carbides and as wear-resistant thin-films. Often Ti(C,N) coatings are multilayered,
i.e. deposited with increasing carbon content in the direction of the coating surface.
By integrating carbon atoms in place of nitrogen atoms into the titanium nitride
crystal lattice, a considerable increase in hardness can be realized, which is positive
for wear resistance but also increases brittleness [Sato78, Schi74, Berg90]. To com-
pensate for this increase in brittleness, Ti(C,N) coatings are deposited as multilayers,
so that residual stresses between the individual coating layers can be reduced.
Ti(C,N) coatings are suited for machining steels with high tensile strength and
thus for higher cutting temperatures.
4.4.2.4 Titanium Aluminium Nitride Coatings ((Ti,Al)N)
The (Ti,Al)N coating system was developed in order to improve the oxidation
resistance, hot hardness and wear-protection properties above the levels of previ-
ously used coatings [Quin87, Knot87]. In comparison to TiN and Ti(C,N) coatings,
(Ti,Al)N coatings have the highest oxidation resistance with a comparably high level
of hardness. Since Ti(Al,N) is a metastable coating system, it can only be deposited
with the PVD, PA-CVD or MT-CVD process.
(Ti,Al)N coatings are a further development of TiN, whereby titanium is substi-
tuted by aluminium by 20–60 at.-%. Depending on the composition, these layers
range from brown (lower Al-content) to black-violet (higher Al-content) [VDI
3824].
Due to their high level of oxidation resistance and hot hardness, the preferred
areas of application are dry machining, hard machining and HSC machining. The
excellent wear resistance of (Ti,Al)N coatings is explained by the fact that, as
opposed to stable compositephases, metastable layers decompose into stable bound-
ary phases or form stable oxides in an oxidizing atmosphere as long as the energy
required to convert into an equilibrium condition is added (e.g. as heat), as it is
the case in dry machining or high-speed machining. In this context, the high oxi-
dation resistance of (Ti,Al)N is derived from the fact that a thin aluminium oxide
layer is formed on the coating surface which is constantly renewed during the cut-
ting process, thus decelerating the progress of wear. With increasing Al-content, the
oxidation resistance of the (Ti,Al)N coating is increased. Despite its much further
improved oxidation resistance compared to TiN and Ti(C,N) coatings, this coat-
ing also fails beyond approx. 800◦C. TiAlN coatings are deposited as monolayer,
multilayer or gradient layers [Lemm03, Leye04, VDI 3824].
(Ti,Al)N is one of the most commonly used high-performance coating systems.
Important measures being taken to further augment the cutting edge durability of
(Ti,Al)N coatings include the development of nanolayers and increasing the con-
tent of aluminium. The primary goal of these developments is to improve their hot
4.4 Coatings 155
hardness, wear resistance and oxidation resistance. In order to differentiate them
from conventional (Ti,Al)N coatings, those with over 50% aluminium content are
designated as (Al,Ti)N coatings.
4.4.2.5 Al-Cr-N Coatings
Besides increasing the aluminium content another approach to further improving the
oxidation resistance and high-temperature properties of (Ti,Al)N coating systems is
adding small amounts of oxide formers such as Cr, Y or Si. One example for this
direction in coating development is the Al-Cr-N coating system. In comparison to
the conventional (Ti,Al)N coating, the (Al,Cr)N coating system has proved to have
a higher resistance to abrasive wear as well as higher hot hardness and oxidation
resistance [Gey04]. Its areas of application include the dry, HSC and hard machining
of steel and non-ferrous metals by turning, drilling and milling [Gey04, Denk04].
4.4.2.6 Aluminium Oxide Coatings (Al2O3)
The exceptional resistance of Al2O3 against both abrasion and diffusion wear and its
simultaneous insensitivity to oxidative wear, known from its use in cutting ceramics,
has made Al2O3 an obvious candidate as a hard material for coating [Sche88].
Because it is very brittle, Al2O3 is generally not used in monolayer form, but
only in combination with other hard materials in multilayer coatings. The electri-
cally nonconductive Al2O3 layers could thusfar not be fabricated with the PVD, but
only with the CVD process for technical reasons. The deposition of the thermody-
namically stable high-temperature modification of aluminium oxide – α-Al2O3 –
by means of CVD process has been established industrially for over two decades.
Due to its excellent wear protection and high performance potential, attempts are
being made to produce Al2O3 layers with the PVD process as well. A first step in
this direction was the deposition of amorphous aluminium oxide (α-Al2O3) with the
help of RF sputtering technology. Because of the low coating rates and consequently
long process durations, this technique is proven to be uneconomical for coating cut-
ting tools. With the help of the pulsed magnetron sputtering process, it has become
possible to coat cutting tools with crystalline γ-Al2O3 layers at substrate temper-
atures of 500–600◦C. γ-Al2O3 can be applied either as a monolayer or multilayer
coating in combination with carbidic or nitridic intermediate layers [Hauz05].
4.4.2.7 Amorphous Carbon Coatings
The concept “amorphous carbon coatings” includes a number of coatings and
coating systems. They are often referred to as “diamond-like carbon” or “DLC”
coatings. These are carbon-based, highly cross-linked amorphous layers with dif-
ferent amounts of sp2 (graphite bond) and sp3 bonds (diamond bonds) as well
as various amounts of embedded hydrogen. In the case of Me-C:H layers, finely
distributed metal carbides are also embedded. DLC coatings are manufactured by
means of plasma-activated PVD and CVD techniques and deposited at coating
temperatures ranging from room temperature to about 300◦C. Their typical layer
156 4 Cutting Tool Materials and Tools
MeC:H a-C:H a-C Diamond
PVD CVD PVDProduction process
800 to 1800 1500 to 3500 3000 to 7000
Coating thickness in µm 1 to 10 1 to 5 1 to 3
Microhardness in HV 0,05
 
1)
0.1 to 1.5 1 to 3 2 to 6Internal stresses in GPa
350 400 450Graphitisation temperaturein °C 2)
+ +++ ++++
Resistance to abrasion
+ +++ +++
Protection of bais material
against corrosion 3)
10000
Resistance to wear due to adhesion
(against steel)
CVD
3 to 10
–
> 600
++++
+++
1) For the microhardness, mean values are indicated. They are obtained from reported measured
values resulting from different compositions, coating thicknesses and internal stresses.
2) The graphitising temperature is the temperature at which amorphous carbon begins to convert
from a three-dimensional lattice into graphite, considerably affecting the characteristics
of the coating.
3) As the aforementioned hard coatings themselves do not essentially corrode (except for MeC),
they protect the basis material from corrosion. (Leaks in hard coatings may lead to the
development of local galvanic elements and to pitting corrosion.)
+++ +++ +++
(+++)
with good
cooling 
Fig. 4.45 Quanties Characteristing the coating and the coating performance, acc. to VDI 3824/1
thicknesses are 1–5 μm. Figure 4.45 shows some characteristic values for the layer
properties of the most important carbon-based hard material layers producible with
PVD and CVD available today [VDI 3824, VDI 2840].
In order to improve their adhesive strength, carbon layers are as a rule deposited
on a hard material layer (e.g. CrN) as a top layer. Graded intermediate layers made
of CrCN have proved favourable for this, in which case hardness and the Young’s
modulus can be adjusted to the mechanical properties of the carbon top layer by
successive exchange of nitrogen and carbon. This procedure makes it possible to
obtain a largely continuous transition from the hard material layer to the carbon
layer in order to optimize the bond between both coating systems [Leye04].
Amorphous carbon layers are characterized by extremely low friction coeffi-
cients in the case of solid body friction, high wear resistance and low tendencies
to adhesion. In cutting technology, these coatings are used above all on tools for
dry machining NE metals (e.g. aluminium or magnesium) and in fine machin-
ing with great success. Further areas of application are in machining graphite and
fibre-reinforced plastics.
4.4 Coatings 157
The property profile of amorphous carbon layers can be purposefully influenced
by process control and by integrating additional chemical elements. This poten-
tial has led to a number of diverse types of carbon coatings in the last several
years. In the following, the essential characteristics and properties of some selected
amorphous carbon layers will be described in close detail. The classification,
designation and characterization of the coatings are derived from VDI guideline
2840.
According to VDI guideline 2840, amorphous carbon coatings can be subdivided
into two groups, hydrogen-free and hydrogenous coatings. Since all carbon layers
contain a certain amount of hydrogen even without adding hydrogen gas (e.g. from
residual gasses), a limit of about 3 at.-% hydrogen is seen as the transition from
hydrogen-free to hydrogenous carbon coatings. With increasing amounts of hydro-
gen, the amount of interconnection between the carbon atoms decreases, leading to
softer layers [VDI 3824, VDI 2840].
As wear-protection coatings on cutting tools, hydrogen-free carbon layers are
used above all in the form of theamorphous carbon layer a-C and the tertraedic
amorphous carbon layer ta-C. The atoms are arranged randomly in amorphous solid
bodies. The only bonds are those between a few isolated atoms. If the sp2 bond
is predominant in the amorphous carbon layer, it is softer, while it is harder if the
amount of sp3 is higher. Depending on the deposition energy, the carbon atoms in the
layer are predominantly arranged in one of the two hybrid states. At low deposition
energies, the amount of sp2 bonds is higher, making the layers softer. These have
the abbreviation “a-C”. With high deposition energies, sp3 hybridizations with a
tetraedic arrangement predominate. The layers exhibit a high hardness level, and the
compressive residual stresses are increased. Tetraedic amorphous carbon coatings
have the designation “ta-C” [VDI 3824, VDI 2840].
Of all amorphous carbon coatings containing hydrogen, it is above all the non-
modified layers that besides carbon contain only hydrogen, designated with “a-C:H”
as well as modified carbon layers that are of any importance. The latter contain
further elements in addition to hydrogen and are classified in accordance with VDI
2840 into the group of carbon layers containing metal and the group of hydrogenous
amorphous carbon layers modified with non-metals [VDI 3824, VDI 2840].
The a-C:H coating system represent the origin of coating with amorphous hydro-
carbon layers. In addition to carbon, it contains 10–30% hydrogen. The latter
is derived from hydrocarbon gases like acetylene, which are employed in the
fabrication process.
If, during vacuum deposition, metals are additionally integrated, layers con-
taining metal are produced. As opposed to a-C:H layers, these are electrically
conductive and can be manufactured with a technically simpler and cost-efficient
DC method. Of all amorphous carbon layers, these therefore have the broadest and
deepest range of application. Amorphous carbon layers containing metal are also
generally known as “a-C:Me” or “a-C:H:Me” (the abbreviation Me-C:H is also
common), where “Me” stands for “metal”. Instead of the abbreviation “Me”, the
integrated metals can also be named specifically, e.g. tungsten (a-C:W, a-C:H:W)
oder titanium (a-C:Ti, a-C:H:Ti). The added metals form finely distributed carbides
158 4 Cutting Tool Materials and Tools
with the carbon in the matrix. By integrating metal carbides, layer adhesion can be
improved and the tribological properties of the coating can be affected [VDI 3824,
VDI 2840].
Hydrogenous amorphous carbon layers modified with non-metals contain non-
metallic elements such as silicon (Si), oxygen (O), nitrogen (N), fluorine (F) or
boron (B), which partially also form carbides. By incorporating different non-
metals, further improvements become possible. Silicon for example helps to
increase temperature resistance. Moreover, different elements can be built into the
layer at the same time. In this way, special layer properties like surface energy
(adhesion tendency, wettability) can be altered.
4.4.2.8 Self-Lubricating Coatings
Self-lubricating coatings include coating systems which include graphite or molyb-
denum disulfide (MoS2). On cutting tools, graphite and MoS2 coatings are applied
as top layers on a supporting hard material layer, where they act as a solid lubricant.
Graphite is deposited in combination with WC as a WC/C multilayer coating system
consisting of several layers of WC and graphite.
Graphite and MoS2 have similar crystal structures. It is characteristic of graphite
that its C atoms are arranged in a plane resulting in a very high level of internal bond
strength due to the small inter-atomic distance. There is a relatively large distance
between the atom layers however, so the layers of atoms can be easily shifted against
each other under the influence of external forces. MoS2 also has a lamellar structure.
There is a very strong chemical bond between the Mo and S atoms forming each
lamella. The bond between two neighbouring lamellae is made between the sulphur
atoms. Due to the intermolecular van der Waals bonds between the sulphur atoms
of two lamellae, these can be shifted towards one another very easily. These slip
planes are responsible for the soft, lubricating effect of graphite and molybdenum
sulphide.
Both coating systems are used primarily for machining materials that tend to
adhere strongly with the tool, for example in the case of dry machining aluminium
wrought alloys. In the case of drilling, these layers decrease the friction between the
chip and the tool because of their “lubricating” effect and thus help to improve chip
removal. In the case of tapping, in which case the machining results are determined
above all by crushing, friction and adhesion processes, these coating systems can
bring about significant improvements in performance even without hard material
intermediate layers.
Because of their properties, both coating systems are gradually worn out during
the cutting process (loss lubrication). Even if the layer appears optically to be worn,
its effect is still partially demonstrable. This is explained by the fact that the coating
material still exists in microscopically small recesses on the tool surface where it can
still be effective. After a complete loss of the soft top layer, the further wear proper-
ties of the tools is then determined by the properties of the supporting hard material
system.
4.4 Coatings 159
4.4.2.9 Diamond Coatings (Crystalline Carbon Coatings)
Coating tools with polycrystalline diamond has been available only since 1990 and
is as such the most recent technique in diamond tool production. Diamond coat-
ings are applied to tools made of cemented carbide or ceramics in a low-pressure
diamond synthesis (a CVD process). The diamond layer consists of pure diamonds
and does not have a binder phase. Industrial deposition temperatures are around
600–1000◦C. This coating technology makes it possible to manufacture even tools
with complex geometries, such as curved surfaces, economically out of diamond
[Leye95]. Application examples for this are drilling tools, end milling cutters and
indexable inserts with chip form geometries. These tools are used in the machining
of plastics, cemented carbide and ceramic green bodies, non-ferrous metals as well
as in the machining of abrasive materials. Diamond-coated tools can not be used to
machine steel and other ferrous materials since the diamond coating is easily worn
due to the solubility of carbon in iron.
We differentiate between thin diamond layers (sd = 1–40 μm) and thick diamond
layers (sd = 0.3–2 mm). Thin diamond layers are used to coat components (tools, for
example) directly. On the other hand, thick diamond layers are generally deposited
on an auxiliary substrate and then detached from it again. Diamond plates produced
in such manner are then used as free-standing diamond (e.g. as radiation windows)
or mounted on supports, usually by means of vacuum soldering in order to produce
tools or other components. Since this can no longer be referred to explicitly as a
layer, we often no longer refer to it as a CVD diamond layer but as a CVD diamond
[VDI 2840].
Several processes have been established for fabricating diamond coatings. The
most important include the hot filament process (HF-CVD), the microwave process
and the plasma jet process (DC arcjet). What all of these processes have in com-
mon is the decomposition of carbonic gases (e.g. methane, acetylene or carbon
monoxide) followed by a deposition of carbon on the substrate surface. The forma-
tion of unwanted graphite must be avoided in the process. Atomic hydrogen plays
a key role in the germ formation and growth of the diamond layers. In the case
of the hot filament method, the best-known process used for depositing CVD thin
diamond layers, the atomic hydrogen required for the process is produced by elec-
tricallyheated filaments made of refractory metals (Ta, W or Re). Gas temperatures
of over 2500◦C are necessary to initiate the required chemical reactions [Lemm04,
VDI 2840].
The rim zone is of particular importance when coating cemented carbides.
Cobalt, which forms the binder phase of the cemented carbide, reduces the germ
formation of diamond on the substrate surface, accelerates graphite formation, and
affects crystal growth and layer adhesion. Cemented carbide substrates must there-
fore be subjected to pretreatment prior to coating, in which case we differentiate
between physical and chemical pretreatment and the introduction of an intermedi-
ate layer. The common process presently is a multi-stage pretreatment consisting of
irradiation, chemical etching, purifying and seeding. The substrate is first purified
and homogenized with microbeam treatment. In the subsequent chemical etching
160 4 Cutting Tool Materials and Tools
process, the cobalt binder phase is removed from the rim zone of the cemented car-
bide substrate. The depth to which the binder phase is removed depends on the back
diffusion of cobalt during the coating process. Due to the high substrate temperature
of 900◦C and long processing times, cobalt diffuses through the etched rim zone to
the substrate surface during the diamond coating process. If the etching depth is
too small, the cobalt that has diffused to the surface reacts with the diamond layer
and leads to the separation of diamonds on the boundary surface. This lowers the
adhesive strength of the diamond coating. If the etching depth is too large, the car-
bide bonds in the binder phase are weakened, also lowering layer adhesion. The
rough surface caused by etching serves mechanically to clamp the diamond layer to
the cemented carbide substrate. Due to the extreme differences between the coating
and substrate material (Young’s modulus and hardness) this mechanical clamping
is of special importance with respect to the adhesive strength of the diamond layer.
After etching, the substrates are again purified and seeded. In order that the diamond
layer will develop, there must be diamond germs on the substrate surface that serve
as starting crystals and from which the layer grows. Seeding can be performed, for
example, by applying small diamond crystals with a diamond suspension [Gram04,
VDI 2840, Lemm04].
Standard diamond layers have a microcrystalline structure (Fig. 4.46). The
crystallites grow from the germs with varying speed depending on their crystal ori-
entation. Slowly growing crystallites are covered by the faster-growing ones. With
increasing layer thickness, the crystallites forming the surface become larger and
larger, resulting in the typical appearance of a microcrystalline CVD diamond layer
with sharp-edged crystal surfaces [VDI 2840].
If growth conditions are set such that new germs are constantly formed and large-
scale growth of individual crystallites becomes impossible, crystallite size remains
in the range of nanometres within the entire layer (Figs. 4.46 and 4.47).This is how
“nanocrystalline” CVD layers are formed, which are much smoother compared to
crystalline layers [VDI 2840].
Both layer types can also be combined in multilayer structures, uniting the pos-
itive attributes of microcrystalline and nanocrystalline diamond layers [Uhlm05]
(Figs. 4.46 and 4.47). This increases the fracture toughness of the entire system and
contains the development of cracks, such that they can no longer easily reach the
Microcrystalline (standard) Nanocrystalline (smooth) Multilayer (smooth)
Microcrystallin
Nanocrystalline
Fig. 4.46 Growth model of diamond coatings, acc. to VDI 2840 and LEMMER [Lemm04]
4.5 Ceramic Cutting Tool Materials 161
Microcrystalline (standard) Nanocrystalline (smooth) Multilayer (smooth)
Fig. 4.47 SEM photographs of CVD diamond coatings on a cemented carbide substrate [VDI
2840, Lemm04]
interface between the first diamond layer and the substrate, potentially causing the
entire surface to chip [Lemm04].
4.4.3 Substrate Pretreatment
Both CVD and particularly PVD coating processes place high demands on the sur-
face conditions of the components that are to be coated. Even the smallest impurities
on the surface of the substrate – be it residual cutting fluid or grease, water deposits,
grinding and polishing agents, hardening salt residues or oxide deposits – can degas
during the coating process, negatively influence that process and lead to coating
adhesion problems. The pretreatment of the substrate therefore has a key role in
creating high and reproducible coating quality [Lugs03, VDI3824].
In substrate pretreatment a distinction is drawn between mechanical pretreatment
(grinding, deburring, microbeaming and polishing) and physicochemical pretreat-
ment (coating-removal, ultrasound-supported purification). In order to optimize
layer adhesion, the parts are subjected to a multi-staged ultrasound purification pro-
cess before charging. In the coating plant, the parts are then first heated in a vacuum,
whereby further impurities can evaporate. Then the tools are etched by means of
argon ion bombardment in order to produce a purely metallic surface. Coating fol-
lows directly thereafter. If the substrates still need to be stored temporarily before
charging, this takes place in a drying closet in order to protect the substrate sur-
faces from water deposits (from air moisture), rust film formation and skin contact
[Lugs03, VDI3824, Gey03].
4.5 Ceramic Cutting Tool Materials
Ceramic materials include all non-metallic, inorganic solid materials. Essentially
these are chemical compounds of metals with non-metallic elements of group III A
to group VII A of the periodic system of elements. A distinction is made between
oxidic and non-oxidic ceramics. The largest group of ceramics are the oxides.
Non-oxidic ceramics are carbides, borides, nitrides and silicons. Some authors dif-
ferentiate again between metallic hard materials (compounds between C, B, N or
162 4 Cutting Tool Materials and Tools
S and Ti, Zr, Nb, Ta, W among others) and non-metallic hard materials such as
diamond, SiC, Si3N4, B4C and BN. Characteristic properties of ceramic materials
are compressive strength, high chemical resistance and high melting temperatures,
which can be derived from the strong covalent and ionic bonds of the atoms
[Salm83a, Horn06].
The following will describe more closely those ceramic materials that are used
as cutting tool materials in cutting technology. Following the usage common in
practice, we will differentiate between cutting ceramics (oxidic and non-oxidic
ceramics) and superhard non-metallic cutting tool materials (diamond and boron
nitride).
4.5.1 Cutting Ceramics
Ceramics have applications in many areas of machining with geometrically defined
cutting part geometries. Mentionable among these are the high-speed machining of
cast iron materials, hard machining or nickel-based alloy machining. The use of
ceramics in the mass production of brake disks, flywheels and like workpieces has
become a matter of course. In the field of automotive technology, silicon nitride
ceramics have been able to capture an important market share in the high-speed
cutting of cast iron workpieces [Schn99]. Hard machining is a primary applica-
tion area of mixed ceramics. “Whisker-strengthened” ceramics are often used in the
area of engine construction for machining engine components made of nickel-based
alloys. Whiskers are needle-shaped monocrystals with a small degree of misorien-
tation in the lattice. Correspondingly, they have high mechanical strength (Rm up
to 7000 N/mm2). Their length is about 20–30 μm, their diameter between 0.1 and
1 μm.
The increased interest in ceramics in production is based not only on the excel-
lent wear properties of ceramic cutting tool materials but most importantly on their
toughnessproperties, which have been clearly increased in recent times. The brit-
tle fracture attributes characteristic of ceramic materials, the wide variation of their
strength properties and the resulting stochastically appearing tool fractures are as
before the main causes limiting the use of these cutting tool materials in machining
technology (in contrast to cemented carbides for example).
New developments and improvements in ceramics have been concentrated there-
fore on further augmentation of the toughness and dependability of these cutting
tool materials in production.
In the case of pure oxidic ceramics, research is focused on improving these prop-
erties by increasing their zirconium oxide content, improving the distribution of
these phases and by making the microstructure more consistent. In the case of mixed
ceramics, finer-grain hard materials are used and titanium carbide is being par-
tially substituted with titanium carbon nitride. Further measures include fibre and/or
whisker reinforcement, reduction of the amount of glass phase at the grain bound-
aries of non-oxidic ceramics as well as the formation of special microstructures
[Kola86, Clau77, Schn99].
4.5 Ceramic Cutting Tool Materials 163
Cutting ceramics
Oxidic
ceramics
Non-oxidic
ceramics
Oxide
ceramics
Mixed
ceramics
Whisker-
reinforced
ceramics
Silicon
nitride
ceramics 
Al2O3
Al2O3+ ZrO2
Al2O3 + TiC
Al2O3 +
 ZrO2 + TiC
Al2O3 +
 SiC-whisker
β-Si3N4
β-Si3N4 + hard materials
α-Sialon
(α + β)-Sialon
Fig. 4.48 Classification of cutting ceramics
Ceramic cutting tool materials can be subdivided into oxidic and non-oxidic
cutting ceramics (Fig. 4.48).
Oxidic cutting ceramics include all cutting tool materials based on aluminium
oxide (Al2O3). A distinction is made between oxide ceramics that only contain
oxides (e.g. ZrO2) in addition to Al2O3, mixed ceramics, which in addition to Al2O3
also contains metallic hard materials (TiC/TiCN) and whisker-reinforced ceramics,
in which SiC whiskers are integrated in the Al2O3 matrix.
Figure 4.49 shows images of fractures of different cutting ceramics taken with a
scanning electron microscope. It is typical of oxidic cutting ceramics that they do
not have a visible binder phase such as cemented carbides (cobalt) have. One can
recognize the globular grains of oxide ceramics, the extremely fine grains of mixed
and whisker-reinforced ceramics as well as the SiC whiskers, which are only a few
μm large in the detail magnification. Characteristic of non-oxidic ceramics based
on Si3N4 is the needle-shaped formation of the crystals. In order to improve wear
resistance, Si3N4 cutting ceramics can be coated with Al2O3 or with multilayer
coating systems made of Al2O3 and TiN [Schn99].
Figure 4.50 shows some physical properties of commercially available ceramic
cutting tool materials. Depending on their chemical composition, these ceramic
types can have considerably diverse properties.
4.5.1.1 Cutting Ceramics Based on Al2O3
Oxide Ceramics
White oxide ceramics are the traditional type of cutting ceramics. Ceramics based
Al2O3 on were already introduced as cutting tool materials at the end of the 1930s
[Kola86]. Inserts were made for a long time out of pure aluminium oxide. Due
164 4 Cutting Tool Materials and Tools
Oxide ceramic
(Al2O3 + ZrO2)
Mixed ceramic
(Al2O3 + TiC )
Silicon nitride ceramic
(Si3N4 + MgO )
Coated silicon nitride ceramic
(Si3N4 + Al2O3)
Detail magnification
(Al2O3 + SiC-Whisker )
Whisker-reinforced ceramic
(Al2O3 + SiC-Whisker )
Fig. 4.49 Fractures of cutting ceramics
Properties
Oxide
ceramic
Al2O3 
Whisker-reinf.
Oxide
ceramic
Al2O3 
Bending
strength 
Compressive
strength
Young‘s
modulus
Vicker
hardness
ρDensity
− -
bB
σdB
+3.5%
ZrO2
+15%
ZrO2
+15%ZrO2
+20%SiC-
Whisker
4.0
1730
700
5000
380
4.2
1750
800
4700
410
3.7
1900
900
-
390E
Mixed ceramic
Al2O3 
Silicon nitride
ceramic
Si3N4 
+10%ZrO2
+5%TiC
+30%
Ti(C,N) +10%Y2O3
4.1
1730
650
4800
390
4.3
1930
620
4800
400
3.3
1750
800
-
g/cm3
N/mm2
N/mm2
103⋅N/mm2
α
Fracture
toughness
Thermal
conductivity
Thermal
expansion 
Melting
point 
λ
T
4.5
16.4
8
-
5.1
15
8
-
8.0
32
-
-
4.2
14.7
8
-
4.5
20
8
-
7.0
3.4
-
N⋅m1/2/mm2
W/m⋅K
10–6⋅K–1
°C
KIC
σ
Fig. 4.50 Physical and mechanical properties of different cutting ceramics and their main
components
4.5 Ceramic Cutting Tool Materials 165
to their brittleness and vulnerability to fracture, such inserts are no longer used
in machining. The pure ceramics used today are dispersion material containing
Al2O3 as well as about 3–15% of finely distributed zirconium dioxide to improve
toughness.
The toughness-enhancing effect of dispersed ZrO2 particles in an Al2O3 matrix
is based on the phase transformation of zirconium dioxide. ZrO2, which exists in
the form of a tetragonal lattice modification in the sintering temperature range
(1400–1600◦C), transforms during cooling into its monoclinic low-temperature
modification. The temperatures in which this transformation takes place depend on
the size of the particles. The smaller the ZrO2 particles are, the lower the transforma-
tion temperature. Since the transformation from the tetragonal to the monoclinical
modification is associated with a volume expansion, various specific mechanisms
of action can assert themselves depending on the size of the particles. The com-
mon effect of all these mechanisms is that they ultimately absorb fracture energy.
The speed of crack development is reduced by microcracking, crack branching, the
stress-induced transformation of small ZrO2 particles as well as crack diversion. The
result of this is that critical cracks only develop at a higher level of energy, which
corresponds to an increase in fracture resistance and an improvement of ductility
[Clau77, Clau84, Zieg86].
Mixed Ceramics
Mixed ceramics (black ceramics) are dispersion materials based on Al2O3 that con-
tain between 5 and 40% of non-oxidic components in the form of TiC or TiCN.
The hard materials in the matrix form finely distributed phases, which limit the
growth of aluminium oxide grains. Correspondingly, these ceramics have a very
fine-grained structure, improved toughness properties and a high level of edge
strength and wear resistance. Compared to pure ceramics, they are harder and have
more favourable thermoshock properties due to their high level of thermal con-
ductivity (Fig. 4.50). The toughness of these ceramics can be further improved by
adding ZrO2.
The development of mixed ceramics is progressing towards finer and finer-
grained cutting tool materials with extremely homogeneous textural structures. As
opposed to conventional cutting ceramics based on aluminium oxide and titanium
carbon nitride with an average grain size of < 2 μm, the more advanced mixed
ceramics have a submicron structure with grain sizes of < 1 μm. The finer-grained
structure increases hardness and bending strength and consequentially the mechan-
ical and thermal loadability, wear resistance and edge strength of ceramic inserts.
Fine-grained mixed ceramics are used in hard-fine machining, e.g. for machin-
ing hardened rolling bearing steels, for hard-fine turning case-hardened automotive
components such as drive wheels, crown wheels, gearwheels or sliding sleeves with
a Rockwell hardness of 54–62 HRC but also for planing and fine planing cast iron
at very high speeds. In the case of hard turning, submicron mixed ceramics are
competing with PCBN cutting tool materials in many areas of application as more
economical alternatives due to their good cost-benefit ratio [Krel97, Schn99].
166 4 CuttingTool Materials and Tools
Whisker-Reinforced Cutting Ceramics
Whisker-reinforced cutting ceramics are cutting tool materials based on Al2O3 with
about 20–40% silicon carbide whiskers. The goal of whisker-reinforcement is to
enhance the toughness properties of ceramic cutting tool materials. The increase
in toughness gained by incorporating whiskers in oxidic ceramics is remarkable.
Compared to mixed ceramics, whisker-reinforced types have up to 60% higher
fracture toughness. The whiskers bring about a more consistent distribution of
mechanical loads in the cutting tool material as well as a more rapid transport of
heat from the thermally highly loaded cutting areas due to their improved heat con-
ductivity. This results in improved fatigue limits and thermoshock resistance, so that
whisker-reinforced cutting ceramics can also be used in wet cut processes.
Fabrication of Oxidic Cutting Ceramics
The starting materials of cutting ceramics, including pressing and sintering auxil-
iaries, are metered according to an exact formula, homogenized in vibration mills
and converted to a compressible powder with a high bulk material stability in a
spray-dryer.
Oxide ceramics, as well as mixed ceramics with small amounts of hard materials,
are compressed at room temperature to inserts (cold pressing) and then sintered.
The sintering temperature is about 1600◦C. Mixed ceramics with a high percent-
age of hard material (> 10%) as well as whisker-reinforced ceramics have to be
hot-pressed, i.e. pressing and sintering take place in one cycle. Due to the special
equipment required for this (e.g. graphite moulds instead of cemented carbide ones),
hot pressing is more cost-intensive than normal sintering. Furthermore, the number
of manufacturable tool geometries is much more limited. Hot isostatic pressing can
be used to subsequently compact sintered ceramic parts, thus reducing the poros-
ity of the cutting tool material. One method variant is hot isostatic compacting of
pre-sintered closed-pore mouldings in an installation.
After sintering/hot pressing, the contact face, and in the case of precision inserts,
the lateral surface and cutting edge chamfer are ground with diamond grinding discs
[Kola86, Grew85].
Properties of Oxide Ceramic Cutting Tool Materials
Because of their high hot hardness and chemical resistance, ceramic cutting tool
materials based on Al2O3 are characterized by excellent wear properties. However,
the highly favourable wear properties of oxide ceramic cutting tool materials is
accompanied by their sensitivity to tensile, bending, impact and thermal shock
stresses. As mentioned at the beginning of this chapter, current further develop-
ment of oxide ceramic cutting tool materials has for this reason been concentrated
on improving toughness and thermoshock properties without diminishing wear
resistance.
4.5 Ceramic Cutting Tool Materials 167
The excellent wear resistance of oxidic cutting tool materials, particularly at high
temperatures, is based among other things on their superior hardness and compres-
sive strength compared to other cutting tool materials at high temperatures. For
example, oxide ceramic cutting tool materials at 1000◦C are still harder than high
speed steel at room temperature (Fig. 4.51).
While the compressive strength of Al2O3 at room temperature roughly corre-
sponds to that of cemented carbide, it is at 1100◦C still as great as that of steel at
room temperature, while neither steel nor cemented carbide are resilient to pres-
sure at 1100◦C. The high compressive strength of ceramics has been a primary
reason for their being privileged as cutting tool materials. But because their band-
ing fracture strength is relatively low, the use of oxide ceramics as the cutting tool
material requires that the process kinematics be selected such that the cutting forces
are effective as much as possible only in the form of compressive stresses [Kola86,
Kola86a, Zieg86].
A further property which makes the use of ceramic cutting tool materials advan-
tageous at high cutting speeds is Al2O3’s low level of creeping. In the case of
cemented carbides, high temperature strength is limited by the property-determining
cobalt phase at temperatures of 800–900◦C. At higher temperatures, creeping pro-
cesses are introduced which are smaller by several orders of magnitude in the case
of Al2O3.
Diamond
Boron nitride
Mixed ceramic
Al2O3 + TiC/TiCN
Oxide ceramic
Al2O3
Carbide
HSSTool steel
0
5000
4000
3000
2000
1000
1250750500 1000250
H
ar
dn
es
s
H
V 
10
Temperature T / K
Fig. 4.51 Hot hardness of cutting tool materials (Source: CeramTec)
168 4 Cutting Tool Materials and Tools
The high wear resistance of oxidic cutting tool materials can also be attributed to
the good chemical resistance of Al2O3. Al2O3 is resistant to oxidation at the cutting
temperature used in practice and has only a small amount of affinity to metallic
materials.
The most crucial disadvantage of ceramic cutting tool materials is their brittle-
ness, i.e. lacking capacity to reduce stress peaks by means of plastic deformation.
The cause of this is the low number of sliding systems in the crystal structures
of ceramic materials. This results in high sensitivity to tensile stresses and low
resistance to mechanical and thermal shock compared with metals.
In the case of a sudden mechanical overload due to impulse-like stress, the cutting
tool material is destroyed by brittle fracture. Because of insufficient ductility, cracks
begin to form when the inner material cohesion is exceeded by external stress. If the
crack reaches a critical size, unstable crack growth appears leading to fracture of
the ceramic tool. The stress intensity factor KI is a parameter for judging the stress
condition at the peak of a crack. Failure of the cutting tool material due to fracture
occurs if the stress intensity factor reaches a critical value, crack toughness (fracture
toughness) KIC. In comparison to other materials, the values of KIC are very low for
ceramics (Fig. 4.50).
Another important disadvantage of oxide ceramics are their relatively low resis-
tance to temperature change. In the case of temperature changes of more than 200◦C,
pure Al2O3 is destroyed. This can only be improved by alloying a component of
superior resistance to temperature change.
Due to their sensitivity to thermoshock, coolants should not be used in conjunc-
tion with oxide ceramics when roughing and planing. Coolants should only be used
to temper the workpiece, e.g. due to narrow workpiece tolerances.
Low bending fracture strength and relatively high sensitivity to impact and
temperature change stresses require a slanted entering and leaving cutting path,
which is associated with delayed load of the cutting edge of the tool (Fig. 4.52).
Due to their low edge strength, chamfered edges should stabilize the cutting
edge.
Areas of Application of Oxide Ceramic Cutting Tool Materials
It must be stressed that, due to their extreme brittleness, a careful adjustment of cut-
ting parameters to tool geometry is the prerequisite of a successful use of cutting
ceramics. There is no room for error, since ceramics can be destroyed immediately
in the form of fracture should the marginal conditions be unfavourable. Cutting
ceramics are therefore almost exclusively used in mass production. This is justified
by the reduction of processing times and the high cost of selecting and adjusting
the cutting tool materials. The main area of application of oxide ceramic cutting
tool materials is rough and finish turning of grey cast iron, case-hardened steels and
heat-treated steels. However, pure oxide ceramics are being increasingly replaced by
silicon nitride ceramics, especially in cast iron processing (Fig. 4.53). The increased
popularity of dry machining however is opening up new areas ofapplication to oxide
ceramics in the case of turning extrusion parts with a small machining allowance.
4.5 Ceramic Cutting Tool Materials 169
1. Chamfering
1. Chamfering and
external turning
2. Face turning3. Face turning
2. Chamfering and
internal turning
1.
Clamping
(external) 
Finished part
2.
Clamping
(internal)
Fig. 4.52 Entering and exit tool path when turning with cutting ceramics (Source: Ford)
Fine-grained mixed ceramics are used in hard-fine machining (e.g. for machining
hardened rolling bearing steels) for fine turning case-hardened automotive com-
ponents such as drive wheels, crown wheels, gearwheels or sliding sleeves with
a Rockwell hardness of 54–62 HRC. They are also used for finish planing and
fine planing cast iron at very high speeds. In the case of hard turning, submicron
mixed ceramics are competing with PCBN cutting tool materials in many areas
of application as more economical alternatives due to their good cost-benefit ratio
[Krel97, Schn99].
Both types of cutting ceramics still however remain restricted with respect to the
carbon content of the machined steels. For example, in the case of steels with a C
content of under 0.35% adhesion occurs on the tool as well as chemical reactions
which increase wear and generally make the process uneconomical.
Whisker-reinforced oxide ceramics have been used with great success for turn-
ing high temperature nickel-based alloys (Fig. 4.54). In contrast to otherwise used
cutting tool materials (cemented carbide or HSS), it is possible to increase the
cutting speed by a factor of 10 or more. In the case of rough and finish milling
Inconel 718 with milling head face cutters, cutting speeds of 800–1000 m/min can
be realized with cutting ceramics based on silicon nitride or whisker-reinforced
aluminium oxide. As opposed to machining with cemented carbide tools, this rep-
resents an increase in performance by factors of 25 or more [Momp93, Gers02,
Krie02, Uhlm04].
170 4 Cutting Tool Materials and Tools
Al2O3-ceramic Si3N4-ceramic
vc = 1000 m/min
f = 0.5 mm
ap = 3.0 mm
vc = 600 m/min 
f = 0.35 mm
ap = 3.0 mm 
300
250
200
150
100
50
0
To
ol
 
lif
e 
qu
an
tit
y 
N
 
non-chamfered chamfered non-chamfered chamfered
Fig. 4.53 Influence of cutting ceramics and entering path on the tool life quantity in brake disc
manufacturing (Source: CeramTec)
Determinant tool life wear criteria
Notch wear
Flank wear
VB VBVBN
vc = 400 m/min
f =
 
0.12 mm
ap = 0.30 mm
Inconel 718
Wet cut
To
ol
 li
fe
 tr
av
el
 p
at
h 
L c
 / m
3000
Cemented
carbide
K10/20
Polycrystalline
cubic boron
nitride
Cutting speed vc / (m/min)
600200100 300
1000
500
400
300
200
100
20 5030 40
Criteria of tool life: VB = 0.3 mm, VBN = 0.3 mm
400
2000
SiC-whisker
reinforced
cutting ceramics
and SiAlON
Fig. 4.54 Achievable tool life travel paths and applicable cutting speeds in finish turning Inconel
718 with uncoated cemented carbide, whisker-reinforced cutting ceramic, silicon nitride ceramic
and PCBN
4.5 Ceramic Cutting Tool Materials 171
Chemical reactions and built-up edge formation when machining light metal
alloys make Al2O3 cutting ceramics unsuitable for machining Al, Mg and Ti alloys.
Cutting ceramics composed of thin ceramic tubules represent a new generation
of ceramic cutting tool materials. In the case of these ceramics, first introduced at
EMO 2005, the tool body is formed of small thin tubes consisting of a tough ceramic
material and filled with a hard, wear-resistant ceramic material. The diameter and
thickness of the tubes, as well as their orientation and filling, are variable. The tubes
are arranged in a plane. The tool is formed with a large number of planes built on
top of each other, where the tubes forming the individual planes can be parallel
from plane to plane or can be arranged alternately at an angle of 45◦/90◦ to each
other. In this way, the properties of the cutting tool material, such as toughness and
wear resistance, can be modified and adjusted to the stresses of various machining
tasks. Initial machining experiments in turning high-temperature nickel-based alloys
as well as in machining hardened steel materials demonstrated the promising wear
properties and performance of this new cutting ceramic [Haid05].
4.5.1.2 Non-oxidic Cutting Ceramics
Among the non-oxidic ceramics (carbides, nitrides, borides, silicides. . .), materi-
als based on Si3N4 have been especially successful as cutting tool materials for
machining purposes. Si3N4 cutting ceramics, in contrast to oxidic cutting ceram-
ics, are characterized by increased toughness and improved thermoshock resistance
(Fig. 4.50). Moreover, they have a higher hot hardness and high temperature
strength. In machining processes involving grey cast iron, they make it possible
to use the highest cutting values while maintaining the highest tool life spans and
low failure rates. The high safety level of these cutting tool materials has been an
especially important factor in the acceptance of Si3N4 cutting tool materials by
manufacturers.
The significantly higher fracture resistance compared to oxide and mixed ceram-
ics is based on the needle-like shape of the hexagonal β-Si3N4 crystals as opposed
to the globular Al2O3 grains. The non-directional growth of the needle-shaped crys-
tals leads to a microstructure made up of mechanically interlinked components
which provides the cutting tool materials with excellent strength properties. The low
thermal expansion of Si3N4 ceramics compared with oxide and mixed ceramics is
responsible among other things for its favourable thermoshock properties. However,
in order to press silicon nitride ceramics completely, sintering auxiliaries (Y2O3,
MgO, Al2O3) are required that form a glass phase and fill in the gaps between the
crystals. This glass/binder phase has a negative influence on the high-temperature
properties of silicon nitride ceramics.
Besides sintering auxiliaries, Si3N4 cutting ceramics can also contain other addi-
tives that affect their crystal structure or texture and thus their properties as well.
Corresponding to their chemical composition and crystallographic structure, the
silicon nitride ceramics available today can be subdivided into three groups:
• I: β-Si3N4 + binder phase (Y2O3, MgO, Al2O3)
• II a: α sialon (α–Si3N4 + Al2O3 + AlN) + binder phase (Y2O3)
172 4 Cutting Tool Materials and Tools
• II b: (α + β) sialons (α Si3N4 + Al2O3 + AlN) + additive (e.g. ytterbium)
+ binder phase (Y2O3)
• III: β silicon nitride + hard materials (e.g. TiN, ZrO2, SiC-whiskers) + binder
phase
The cutting tool materials from Groups I and II are fabricated by means of hot
pressing, sintering, hot isostatic pressing or a combination of these methods. Most
of the Si3N4 cutting tool materials available on the market today belong to Group I.
The cutting tool materials from Group II are usually designated as sialons. Silicon
nitride can receive up to 60% aluminium oxide in solid solution. Some nitrogen
atoms are replaced in the process by oxygen atoms and silicon atoms by aluminium
atoms. While the α sialon mixed crystals have a globular form, β sialon mixed crys-
tals are stem-shaped. Including special additives makes it possible to stabilize the
β sialons, thereby creating a cutting tool material the structure of which consists
of both α and β mixed crystals. The amount of additives determines the amount
of β sialons in the cutting tool material’s microstructure. In comparison to the cut-
ting tool materials from Group I, sialons are harder, more chemically resistant and
have an increased resistance to oxidation. The manufacturing method in this case is
sintering, which can sometimes be followed by hot isostaticpressing.
Group III can include Si3N4 cutting tool materials, the properties of which can
be specifically altered by adding hard materials such as titanium nitride, titanium
carbide, zirconium oxide or Sic whiskers.
Areas of Application of Non-oxidic Cutting Ceramics
The classic field of Si3N4 cutting tool materials is grey cast iron machining. In this
case, usually the tougher silicon nitride ceramics of Group I are preferred, espe-
cially in automated manufacture. Due to the high fracture toughness of these cutting
tool materials, large feeds, high cutting speeds and thus large volume removal rates
can be realized when machining cast iron materials in smooth and interrupted cut
[Köni86, Köni87, Schn99]. For example, when turning automotive brake discs,
using Si3N4 ceramics improves tool life quantity considerably compared with oxide
ceramics (Fig. 4.53). Improved resistance to fracture makes it possible, as shown in
Fig. 4.52, to omit workpiece chamfering as long as one forgoes maximal utilization
of the cutting tool material.
The main area of application of sialons is in turning nickel-based alloys under
finishing or average roughing conditions (Fig. 4.54). They are therefore in direct
competition with whisker-reinforced oxide ceramics. Like the latter, sialons can be
used for machining nickel-based alloys at much higher cutting speeds than cemented
carbide tools. They are also good for high-speed milling of nickel-based alloys with
tools mounted with indexable inserts at cutting speeds of 800–1000 m/min.
The wear resistance of silicon nitride ceramics is slightly lower than that of oxide
ceramics. Silicon nitride cutting tool materials exhibit a strong affinity to iron and
oxygen under machining conditions. They wear very quickly in the case of steel
machining, so the use of these materials groups currently remains uneconomical.
4.5 Ceramic Cutting Tool Materials 173
Coating silicon nitride ceramics opens up further prospects for their use. The
classic coating materials are TiN, TiC, TiCN and Al2O3, which are applied in dif-
ferent combinations and layer thicknesses. A multilayer coating system consisting
of aluminium oxide and titanium nitride remains the most popular. The service life
of Si3N4 can be considerably increased by applying a wear and diffusion inhibiting
coating, especially when machining cast iron with globular graphite (e.g. GJV40)
[Schn99]. Coated silicon nitride ceramics are also suitable therefore for cases that
had hitherto been reserved for cemented carbides. As opposed to cemented car-
bides, they allow for higher cutting speeds and thus for shorter machining times and
in many cases also for dry machining [Schn99].
4.5.2 Superhard Non-metallic Cutting Tool Materials
In machining technology, cutting tool materials based on diamond and boron nitride
are designated as superhard and non-metallic. According to the definition, both are
ceramic cutting tool materials, where diamond is assigned to monatomic, boron
nitride to non-oxidic ceramics.
4.5.2.1 Diamond as a Cutting Tool Material
Elemental carbon appears in the two crystal modifications graphite and diamond.
Diamond solidifies in the cubic crystalline lattice system, in which the C atoms are
bonded covalently and tetraedically. The extremely high bond and lattice energy is
the reason why diamond is the hardest of all known materials. Diamond is equally
superior to all other known hard materials with respect to thermal conductivity.
Figure 4.55 shows the hardness and thermal conductivity of some materials that
tend to be used as abrasives as well as different carbides, which are components of
cemented carbide tools and permits a comparison with diamond.
Classification of Diamond Cutting Tool Materials
To classify diamond cutting tool materials, we differentiate between natural and syn-
thetic diamond, which both can appear in monocrystalline or polycrystalline form.
According to DIN ISO 513, the identification letters of polycrystalline diamond
are DP while those of monocrystalline diamond are DM (Fig. 4.3). Following the
usage common in practice and in the literature, we will also refer to polycrystalline
diamond cutting tool materials with the abbreviation PCD.
(a) Natural Diamond
Natural diamond is only of importance for the purposes of cutting with geomet-
rically defined cutting edges in its monocrystalline form. Although diamond
also exists in nature in polycrystalline form (ballas, carbonado), these types of
diamond are of lesser interest since the synthetic fabrication of polycrystalline
diamond is more advantageous both economically as well as technologically.
One significant property of monocrystalline diamond is the anisotropy
(directional dependence) of mechanical properties such as hardness, strength
174 4 Cutting Tool Materials and Tools
Hardness HV Thermal conductivityλRT / (W/m⋅K)
Diamond
cBN
B4C
SiC
TiC
10001002500500010000 10000
α -Al2O3
Fig. 4.55 Hardness and thermal conductivity of diamonds in comparison to other hard materials
or elastic modulus. While this directional dependence is compensated in the
case of polycrystalline materials by the completely random distribution of the
individual crystals, anisotropy remains in monocrystalline materials due to the
aligned orientation of the crystal lattice.
This also the cause of the cleavability of monocrystalline diamonds in four
select cleavage directions. From this we can see that the position of the lattice
directions must be known both for grinding monocrystalline diamond as well as
for using it as a tool. While grinding must always take place in the direction of
lowest hardness, monocrystalline diamond tools have to be oriented in the tool
holder such that the cutting force points in the direction of a hardness maximum.
(b) Synthetic Diamond
The fabrication of synthetic diamonds takes place using a catalyst solution in a
pressure and temperature range in which diamond forms a stable phase. By a
careful selection of pressure and temperature, the crystal growth rate and dimen-
sions of the crystals can be controlled within a range of a few micrometers to
several millimetres and specific physical properties like purity or porosity can
be influenced.
Diamond synthesis results in monocrystalline diamond particles that are
utilized in machining with geometrically defined cutting edges as synthetic
monocrystalline diamond tool bodies or for further processing into polycrys-
talline cutting parts.
Fabrication and Designs of Diamond Tool Bodies
(a) Monocrystalline Diamond Tools
Monocrystalline diamond tool bodies are manufactured with both natural diamond
and synthetic monocrystals [Heus96]. The most common cutting edge shapes of
4.5 Ceramic Cutting Tool Materials 175
Tool with round cutting edgeTool with one cutting edge
κr ≈ 45°
κ'r ≈ 2°
κ'r ≈ 2°
κr ≈ 18°
Tool with facette shape
160°
Fig. 4.56 Shapes of monocrystalline diamond tools
monocrystalline diamond tools are shown in Fig. 4.56 [Wein69]. The tool with one
cutting edge is used both for drilling as well as for external turning, whereby in the
latter case a corner radius is ground onto it in order to improve the surface profile.
By selecting a very small minor cutting edge lead angle κ ′r < 2◦, the minor cutting
edge can be designed as a broad-tool cutting edge which takes on the function of
further polishing the surface.
Tool variants with round cutting edges have the advantage of having a very large
useful cutting edge length, but due to the curvature of the cutting edge they have
unfavourable chip formation conditions and relatively high passive forces. In the
case of tool bodies with facette shapes, three to five cutting edges are ground on,
whereby two neighbouring edges form an angle of about 160◦. The tool is adjusted
so that the minor cutting edge has a very small lead angle as acts