The international journal of advanced manufacturing technology, tập 58, số 5 8, 2012

Int J Adv Manuf Technol (2012) 58:421–429
DOI 10.1007/s00170-011-3420-5

ORIGINAL ARTICLE

Performance enhancements of high-pressure die-casting die
processed by biomimetic laser-remelting process
Zhi-xin Jia & Ji-qiang Li & Li-Jun Liu & Hong Zhou

Received: 16 December 2010 / Accepted: 30 May 2011 / Published online: 9 June 2011
# Springer-Verlag London Limited 2011

Abstract Die service life improvement is an important
problem in high-pressure die-casting industry. Experiment
results on die steel shows that biomimetic laser-remelting
process provides a promising method to improve the
service life of die-casting die. A casting with uneven wall
thickness was selected and problems existing in die-casting
production were analyzed. The corresponding die-casting
die was processed by biomimetic laser-remelting process.
The application result indicates that the service life of the
die processed by biomimetic laser-remelting process has
been increased from 12,000 to 28,000 shots, which is more
than twice that of no processed one under real high-pressure
die-casting conditions. The application of laser-remelting
process provides desirable micro-structural changes in
biomimetic units, which induces the intensified particles
effect for improving the service life.
Keywords Die-casting die . Thermal fatigue .
Laser-remelting process

1 Introduction
Die casting is a high-volume production process, which
produces geometrically complex parts of nonferrous metals
Z.-x. Jia (*) : J.-q. Li : L.-J. Liu
Ningbo Institute of Technology, Zhejiang University,
1# Qianhu Road,
Ningbo 315100, People’s Republic of China
e-mail:
H. Zhou
The Key Lab of Automobile Materials,
The Ministry of Education, Jilin University,
5988# Renmin Road,
Changchun 130025, People’s Republic of China

with excellent surface finishes and low scrap rate. The die
castings are used extensively in automobile, motorcycle,
computer, and consumer electronics. These die castings are
generally produced by using two steel die halves called the
cover-die half and ejector-die half separately. Each of the
die halves usually contains a portion of the die cavity. The
process sequences are: (a) die closing, (b) cavity filling, (c)
casting solidification, (d) parts ejection, and (e) lubrication.
The most important modes of failure in die-casting dies are
thermal cracking, soldering, and corrosion.
Die wear and failure is a significant issue in diecasting industry, owing to the high cost of dies.
Nevertheless, owing to the harshness of service condition
of the die-casting dies, the complexity of thermal fatigue
processes, and the variety of factors affecting the process,
die wear and failure has been a technical difficulty in

die-casting industries for many years. In order to prolong
the service life of die-casting dies, many researchers have
been engaged in the theoretical and experimental studies
related.
Research group from The Ohio State University, USA,
did a lot of work aiming at elucidating the life-limiting
failure mechanisms in the die-casting die through experiments and CAE analysis [1–5]. Venkatesan and Shivpuri [1,
2] carried out experiments under actual production conditions for a range of process and geometrical conditions
with the accelerated erosive wear of core pins being used as
a surrogate measure of die erosive wear. Yu. et al. gave a
study of corrosion of die materials and die coatings in
aluminum die casting [3]. They also studied effects of
molten aluminum on H13 dies and coatings [4]. Their
experiments have shown that single-layer hard PVD and
CVD coatings do not protect the die steel surface from
cracking. Kulkarni et al. investigated the thermal cracking
behavior on nitrided die steels in liquid aluminum process-


422

ing [5]. Srivastava et al. developed a model to predict the
thermal fatigue cracking using FEM software[6].
Persson et al. studied the simulation and evaluation of
thermal fatigue cracking of hot work tool steels [7, 8].
Domkin et al. studied the soldering and did some work to
tackle the problem of die life-time prediction based on a
quantitative analysis of soldering in the framework of the
full 3D simulations of the die casting process [9]. Zheng et
al. established an evaluation system for the surface defect of

casting and introduced artificial neural network to generalize the correlation between surface defects and die-casting
parameters, such as mold temperature, pouring temperature,
and injection velocity [10]. Klobcar et al. analyze the
influence of aluminum alloy die-casting parameters, die
material, and die geometry on in-service tool life by
immersion testing and FEM method [11].
With the development of laser technology, laser
processing method is used to change the property of
die material. Grum et al. reported results of CO2 laser
repair surfacing of maraging steel with a Ni–Co–Mo alloy
similar to the maraging steel [12]. After laser surfacing of
the DIN 1.2799 maraging steel a very favorable throughdepth residual-stress profile of the surfaced layer and the
heat-affected zone is obtained. Compressive residual
stresses in the surface layer reduce the risk of formation
and propagation of surface cracks. Such a state of stress
will considerably extend the tool life. Persson et al.
studied the life-limiting failure mechanisms in dies aimed
for brass die casting [13]. They examined and evaluated
cavity inserts and cores with respect to failure mechanisms
after use in actual brass die casting. They found that the
dominating failure mechanism in the investigated tools
was thermal fatigue cracking.
In order to improve the thermal conductivity of H13
die material, some studies that have been carried out to
develop molds with higher thermal conductivity have
concentrated on mixing copper and steel. Beal et al.
manufacture the 3D structures from a mixture of H13
and copper powders by using a laser beam to sinter or
melt the mixture of H13/Copper powder. The method
employed is based on the layer manufacturing technology [14]. Khalid et al. presents a novel approach to

replace a conventional steel die by a bimetallic die made
of Moldmax copper alloy coated with a protective layer of
steel using laser cladding technology, direct metal deposition on the cavity surface for high-pressure die casting of
aluminum alloys [15].
Nature provides a whole host of superior multifunctional structures that can be used as inspirational systems
for the design and synthesis of new, technologically
important materials and devices. Since the 1980s, Ren et
al. has been dedicating to the study of the cuticle
morphologies and principles of soil animals. They found

Int J Adv Manuf Technol (2012) 58:421–429

that soil animals have “nonsmooth construction units,”
which provide excellent anti-wear properties against soil
[16]. Recent works in the research group of Jilin
University, China, also found that a considerable effect
not only on wear resistance [17], but also on the thermal
fatigue resistance [18, 19] when applied biomimetic
principle on the die and tool surfaces to form a series of
biomimetic units by laser. Experiments were focused on
the effects of laser input energy and biomimetic unit
shape. Zhang et al. studied the size of units and
investigated its effect on thermal fatigue behavior of
3Cr2W8V steel [19]. They also studied the tensile
property of H13 die steel with convex-shaped biomimetic
surface [20]. Shan et al. did some experiments on injection
molds by mimicking the injection conditions. The results
showed that the adhesion biomimetic molds have a
beneficial effect on decreasing the adhesion to eject
polymer parts [21].

Studies [18–20] have shown that the biomimetic surface
with units in varying shapes and distributions has an
enhanced resistance not only to the thermal fatigue crack
initiation but also to the crack propagation. But so far these
studies are merely experiment result in laboratory on
specimen with simplified geometry by mimicking the real
conditions. Regarding the complex-shaped casting, especially the performance of die-casting die processed by
biomimetic laser-remelting process under actual production
conditions is still scarce.
In this paper, a set of die-casting die made of H13 is
chosen to be processed by biomimetic laser-remelting
process and its performance under actual production
conditions is investigated. The real die-casting conditions
are supplied by our partner, Donghao Die-casting Co.,
Ltd. The application result shows that service life of diecasting die processed by biomimetic laser-remelting
process is prolonged from 12,000 to 28,000 shots. The
purpose of this study is to further reveal the effectiveness
of thermal-fatigue-resistant mechanism of the units under
actual production conditions, and finally to lay a
foundation for the application of biomimetic laserremelting process in the design and manufacturing of
die-casting dies in the future.
The rest of the paper is organized as follows. Section 2
gives the requirements and material of the selected casting,
the die-casting parameters, the main problems, and the
service life of the die in die-casting production. Section 3
shows the experiment parameters, method of biomimetic
laser-remelting process, and the performance of the diecasting die under actual production conditions. Section 4
illustrates the microstructure of the unit. Section 5
describes the application of biomimetic laser-remelting
process on the succeeded die-casting die. Section 6 gives

conclusions.


Int J Adv Manuf Technol (2012) 58:421–429

423

2 The die casting and the die-casting die
Due to the high cost of die-casting die, one die casting and
the corresponding die were selected elaborately.
2.1 The characteristics of the selected aluminum die casting
The selected aluminum casting was produced by highpressure die casting, as shown in Fig. 1, called cover, which
is used in vehicles. The material of the casting is ZL102.
Though the geometry of the casting is not very complex, the
dimension accuracy and the surface roughness are required
strictly. The inner surface of die casting is required to keep
the original die-casting surface. There are two platforms
which have flatness checking requirements. The outer
surface of the casting is cleaned by shot blast and then
sprayed with black paint, as depicted in Fig. 1. The average
wall thickness of the die casting is about 7 mm, which is
thicker than general castings. Moreover, the wall thickness is
not even. In the two-platform region, the max thickness
reaches 18 mm, which create areas of high temperatures
during solidification, the so called hot spots. Furthermore,
there are sharp angles or edges near the ribs on outer surface
and platforms on inner surface, which are known to promote
or increase the risk of soldering [9] and corner cracking [6].
The strict requirements, uneven wall thickness, and corners
in small radius lead to great difficulties in die-casting

production and short service life of the die-casting die.
There are two reasons for this casting is selected. One is the
short service life of the die-casting die, which embarrassed our
partner very much. The other is our partner produces the
casting in large quantities, about 10,000 pieces per month for
the customer, which gives the great convenience to investigate
change of the service life of die before and after being
processed by biomimetic laser-remelting process.
2.2 The die-casting process parameters
The processing parameters for the selected die casting are
listed below:
Preheat temperature of the die, 200∼220°C

Temperature of the aluminum molten liquid, 660°C
Die cooling temperature, 250∼300°C
Clamping force, 2,800 MPa
Filling time, 6 s
Inlet temperature of circulating water, 25°C
Outlet temperature of circulating water, 35°C
One cycle time, 55 s
2.3 The defects regions on the casting
In real die-casting production, defects on outer surface of
the casting appear firstly on edges with small radius of the
ribs, as shown in the red circle in Fig. 2a. While defects on
inner surface of the die casting are concentrated on the
boundary edges of the platforms, as shown in the red lines
in Fig. 2b.
2.4 The service life of the die-casting die made
by conventional process
The die-casting die consists of two separate halves: the ejector

die on the “bottom” side of the casting and the cover die on the
“top” side of the casting. The die halves are manufactured of
hardened H13 die steel (0.36% C, 1.09% Si, 0.32% Mn,
5.12% Cr, 1.32% Mo, 0.80% V, and <0.023% P and S), which
is the most commonly used die material.
The manufacturing procedures of our partner for the die
are: (1) rough milling, (2) heat treatment and quenching
operation, (3) EDM machining, and (4) polishing to obtain
low roughness of the die surfaces.
When the die was used to die-casting production, the
existing problems on die castings occurred as follows. After
about 2,500 shots, micro-cracks on the die surfaces with
small radius were generated, which were invisible but could
be touched with unsmooth feeling. While on the casting
surface regions corresponding to micro-cracks, small poxes
were pitted, which increased the roughness of the casting.
Then the two die halves were disassembled. Cleaning and
polishing maintenance works were done to the die surfaces,
which called regular maintenance. After about 5,000 shots,
the cracks on the surface of the die could be observed. So

Fig. 1 Aluminum die casting
(material brand: ZL102)

(a) Outer surface

(b) Inner surface


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Int J Adv Manuf Technol (2012) 58:421–429

Fig. 2 The defect regions of the
die casting

(a) Defect regions on outer surface
the die halves were disassembled and sent to a heat
treatment company to do tempering. Then the die surfaces
were polished and the die halves were reassembled. After
about 7,500 shots, a regular maintenance was done again.
After 10,000 shots, the die was tempered again. The wear,
cracks and soldering on the die surface were even worse,
and more manual revised work needed to be done to the
surface of die castings. As the cracks propagated, then
remedy works by argon-arc welding were needed to repair
the die in order to keep the die in production, which means
that a lot of time are spent on repairing and polishing and
leads to low production efficiency. At last, the die was
abandoned after 12,000 shots and a new one is needed for
production.

3 Experiment method for biomimetic laser-remelting
process

(b) Defect regions on inner surface

The software system provides the tools for user to obtain
a series of points on the die surfaces, then form the route
and generate the NC code. Consequently, the laser

equipment is controlled by user’s NC program to finish
the required route.
3.2 Laser parameters
The laser parameters used for processing the die-casting die
are listed below.
Scanning speed
Electricity
Frequency
Pulse duration
Single-pulse energy
Defocus value

0.5 mm/s
150 A
5 Hz
8 ms
333 J
5 mm

Figure 4a, b shows the pictures in biomimetic laserremelting processing for the cover half and the ejector half
separately.

3.1 Experiment equipment
Biomimetic laser-remelting process is processed on a Laser
Welding Machine, as shown in Fig. 3. The laser equipment
is current-feedback fiber-optic Welder WF300, Han’s Laser.
It is composed of laser generator, worktable, a set of
software system and cooling water tank. The worktable is
moveable along the x and y directions, and the laser is
moveable along z-axis.

3.3 The region determination for the biomimetic
laser-remelting process
As mentioned above in Section 2.2, the die used formerly has
demonstrated the regions where the defects are prone to
generation. The cracked regions are the weakest in structure,

Fig. 3 The laser equipment

Cooling system

Laser

Optical system

Worktable

Industrial computer


Int J Adv Manuf Technol (2012) 58:421–429

425

Fig. 4 The die halves in biomimetic laser-remelting process

(a) The cover half
strength, and thermal resistance. If such regions are to be
strengthened, the service life of the die-casting die will be
prolonged. Hence, the biomimetic laser-remelting process is

carried out on the corresponding regions with the laser
parameters given in Section 3.2. The two die halves processed
by the biomimetic laser-remelting process are shown in Fig. 5.
3.4 The application investigation of die-casting die
processed by biomimetic laser-remelting process
After the die-casting die was processed by laser-remelting
process, it was delivered to our partner. We followed the die
for up to 3 months.
Our partner assembled the die-casting die processed by
laser-remelting process on die-casting machine in operation.
The first ten castings were checked and measured. It was
found that the local thickness of the casting corresponding to
the regions on die surface processed by laser remelting
protruded out about 0.08∼0.09mm. The ejector die and the
cover die were disassembled and manufactured by EDM
separately. Then the processed die was in production
continuously for 18,000 shots. Some unsmoothed spots
appeared on the die castings after 18,000 shots just like that
after 2,500 shots for the unprocessed die. Then the die halves
were disassembled and sent to a heat treatment company to do
tempering. After the die surface was cleaned and polished, it

(b) The ejector half
was put into operation for further 5,000 shots (up to 2,300
shots in total). Then cracks was found near the ejector pin
opposite to the main runner because the ejector pin decreased
the local wall thickness of the ejector half, as shown in Fig. 6.
Industrial experience reported in the past has indicated that
the part exposed to the liquid metal attack in front of the
gates exhibit the highest level washout [1]. In order to

remedy the cracks, the ejector pin hole was blocked and
welded by argon-arc welding. Another 5,000 shots were
obtained and two big cracks were found in the welding
region and some small cracks appeared around the small
radius regions processed by laser. At last the die was
abandoned. So the ultimate service life of the die processed
by biomimetic laser remelting process is 28,000 shots, which
is more than twice that of no processed one. The die-casting
die after 28,000 shots is shown in Fig. 7.

4 Experimental details on specimen and discussion
4.1 The microhardness value of the unit
As the die-casting die is expensive, prior to processing the die
by biomimetic laser-remelting process, a specimen with size of
40×20×3 mm was machined by WEDM from the waste
material in manufacturing the die halves. The surface of the

Fig. 5 The die-casting die after
biomimetic laser-remelting
process

(a) Cover-die half

(b) Ejector-die half


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Fig. 6 The die half after 23,000
shots and the die casting

(a) The die half after 2,3000 shots
specimen was polished using progressively finer grades of
silicon carbide impregnated emery paper, which is as same as
the die surface. Subsequently, the specimens were processed
by the laser equipment with the parameters given in section 3.2.
After the laser processing, a transverse section was
obtained and the standard method of metallography was
followed to prepare the specimen for the microstructure
analysis and the microhardness measurement.
The surface morphology of the specimen was put into
observation under optical microscope. Figure 8 shows the
cross-section appearance of the nonsmooth unit which
involves the bright field surrounded by the parabolashaped contour line. The unit is comprised of the melting
zone and the transitional zone. By the measurement, the
unit width is around 900um, while the depth is 800um.
The microharness was tested on microharness tester DHV1000. The microharness values obtained along x axis and y
axis were listed in Figs. 9 and 10, respectively. We can see
the decrease from the unit hardness to the transition zone and
the substrate hardness. The hardness of the unit varies from
562 to 608 HV, and the transition zone varies from 470 to
500 HV while the substrate microhardness is around 460 HV.
4.2 Microstructure of the unit
The corroded cross-section microstructures of the unit by
biomimetic laser-remelting process were observed by

(b) Die casting

scanning electron microscope, as shown in Fig. 11 (model
HITACHI S-4800, Japan).
Compared with the substrate microstructure, the unit
structure is refined greatly, as shown in Fig. 11. Under the
condition of the laser super fast heating, the larger degree of
overheating was generated to cause the effective driving
force of phase transformation, which contributed to the
large number of the austenite nucleation, Because of the
super fast heating, there was little time for austenite grain to
grow up, finally the refinement of grains and alteration of
microstructure were induced. During the subsequent martensitic transformation, the tiny martensitic structure was
formed. According to Zhou [12], everything happened in
both the super fast heating and cooling induced the
formation of tiny martensitic structure, which contributed
effectively to high hardness and strength. In the transition
area, dentrite structure was formed.
4.3 Discussion on the enhancement of resistance to crack
for the die surface with laser remelting unit
The cracking phenomenon can be divided into two stages
[5]: (1) crack initiation: during the first few cycles, high
thermal gradient can result in shock, causing crack
initiation. (2) crack propagation: once the crack has
initiated, the propagation depends entirely on the toughness of the substrate and the thermal stresses imposed.

Fig. 7 The die-casting die after
28,000 shots

(a) Cover-die half

(b) Ejector-die half

Int J Adv Manuf Technol (2012) 58:421–429

x

427
700

0.00
8

7

6

5

4

3

2

1

605.5

Microhardness (HV)

500

The Matrix
y

Fig. 8 Profile of the unit on the cross section of the specimen

700
586.4

564.8
557.7

480.7

Microhardness (HV)

500
494.9

581.7

570.4

535
486 475.3

471.6

458.7

400

300

200

After a number of thermal cycles, the die becomes so soft
that the applied thermal stresses are enough to cause
plastic deformation. The plastic strains keep on accumulating, resulting in low-cycle fatigue cracking.
According to former works by Zhou [18] and Zhang
[19], when applied laser-remelting process on the die steel
surface, there had a considerable effect not only on
improving the wear resistance and decreasing the adhesion
against parts, but also on inhibiting the initiation and
propagation of thermal fatigue crack.
The refinement of the grains and martensitic high
dislocation density were the primary factors enhancing the
hardness of the units processed by laser. When the thermal
fatigue crack met the units, account for the high hardness in
the units, the cracks could not drill through and finally
stopped near the units. This means that the crack development has been subjected to large obstacles, which forced
the cracks to change the direction. Moreover, the crack
propagation was prolonged, which decreases the rate of
crack propagation.
Hence, analysis mentioned above gives explanation why
the service life of the die processed by biomimetic laserremelting process was improved compared to that without
being processed in real pressure die-casting condition.

562.1

596.4

570.4

Transition Area

600

624.2

608.5

600

The melting zone

497.2
460.5

400

100

0
0.03

0.08 0.15 0.25

0.35

0.45

0.55

0.65 0.75 0.85 0.95 1.05

Distance along y axis (mm)

Fig. 10 The microhardness value of the cell along y direction

5 The succeeded die-casting die processed by biomimetic
laser-remelting process
As the first die-casting processed by biomimetic laserremelting process achieved good results. The succeeded
die-casting die made by H13 was processed again with the
same laser parameters as the former one. However,
considering the sinkage of the region processed by
biomimetic laser-remelting process, some adjustments were
made in the manufacturing of the die.
The manufacturing procedures of the second die are:
(1) rough milling, (2) heat treatment and quenching
operation, (3) roughing EDM machining with allowance
of 0.2mm, (4) biomimetic laser-remelting process operation, (5) finishing EDM machining with removal of
0.2 mm, and (6) polishing to obtain low roughness of the
die surfaces. So the biomimetic laser-remelting process
can be treated as one step in die making. Figure 12
shows the second die-casting die processed by biomimetic
laser-remelting process.
The service life of the second die-casting die processed by biomimetic laser-remelting process is 29,000
shots, which also demonstrated the effectiveness of the

process.

300

6 Conclusions

200
100
0
0

0.25

0.5

0.75

1

1.25

1.5

Distance along x axis (mm)

Fig. 9 The microhardness value of the cell along x direction

1.75

1. The service life of high-pressure die-casting dies can be

improved substantially by biomimetic laser-remelting
process.
2. The units of H13 steel processed by laser-remelting
process have beneficial effects on inhibiting the thermal
fatigue crack initiation and propagation. The biomimetic


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Int J Adv Manuf Technol (2012) 58:421–429

Fig. 11 SEM microstructure at
higher magnification near the
area of the unit. a, b The
microstructure near the area of
the unit. c The microstructure of
transition area. d–f The microstructure of the unit

The Matrix

Transition Area

The Matrix

Transition Area

(a)

(b)

(c)

(d)

(e)

(f)

die-casting die surface possesses superior resistance to
thermal fatigue compared to that without being processed.
So the die-casting die obtains a longer service life.
Fig. 12 The second die-casting
die processed by biomimetic
laser-remelting process

The unit

The unit

3. The biomimetic laser-remelting process provides a
promising way to enhance the performance of diecasting die.


Int J Adv Manuf Technol (2012) 58:421–429

4. Though the result of our experiment under real diecasting condition shows a bright way to resist the crack,
a lot of research works are still needed to reveal the
efficiency of the biomimetic laser-remelting process
further. For castings with the intricate shape, how to
determine and evaluate the layout of the regions for

laser remelting are our future works.
Acknowledgments This research was supported by the Ningbo
Natural Science Foundation (no. 2011A610149) and Zhejiang Natural
Science Foundation (no. Y1100073). The authors wish to thank WH
Yang and YB Zhang from Ningbo Donghao Die-casting Co., Ltd. for
their valuable suggestions and contribution to this research work.

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Tensile property of H13 die steel with convex-shaped biomimetic
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Int J Adv Manuf Technol (2012) 58:431–441
DOI 10.1007/s00170-011-3401-8

ORIGINAL ARTICLE

Implications of the reduction of cutting fluid in drilling AISI
P20 steel with carbide tools
Rodrigo P. Zeilmann & Gerson L. Nicola & Tiago Vacaro &
Cleiton R. Teixeira & Roland Heiler

Received: 17 January 2011 / Accepted: 19 May 2011 / Published online: 3 June 2011
# Springer-Verlag London Limited 2011

Abstract The machining of hardened steel is becoming
increasingly important in manufacturing processes. Machined
parts made with hardened steel are often subjected to high
service demands, which require great resistance and quality.
The machining of this material submits the tools to high
mechanical and thermal loads, which increases the tool wear
and affects the surface integrity of the part. In that context, this
work presents a study of drilling of AISI P20 steel with
carbide tools, analyzing the effects on the process caused by
the reduction of cutting fluid supply and its relation with the
tool wear and the surface integrity of the piece. The major
problem observed in the tests was a difficulty for chips to flow
through the drill flute, compromising their expulsion from the
hole. After a careful analysis, a different machining strategy
was adopted to solve the problem.
Keywords Machining . Drilling process . Cutting tools .
Wear . Surface analysis

R. P. Zeilmann (*) : G. L. Nicola : T. Vacaro
Center of Exact Sciences and Technology,
University of Caxias do Sul (UCS),
1130, Francisco Getúlio Vargas,
95070–560, Caxias do Sul, Brazil
e-mail:
C. R. Teixeira
Federal University of Rio Grande (FURG),
Rio Grande, Brazil
R. Heiler
Hochschule für Technik und Wirtschaft (HTW),

Berlin, Germany

1 Introduction
The machining of hardened steel is a topic of great interest
for industrial production and scientific research. Some
machine components are made of hardened steel materials
and are required to function near their physical limits.
Recent developments in machine tools as well as in process
technology focus on cutting hardened steel and rapidly lead
to an increased awareness of the industrial relevance of
hard cutting [1]. One major problem of machining hardened
steels is the tool wear caused by the hardness of the
material [2]. Although hard machining avoids the shape and
geometrical errors that could occur on a workpiece when
subjected to heat treatment after machining and reduces the
rework, it increases the thermal loads on the tool.
The heat generation and friction between tool and chip
usually limit machining performance in metal-cutting
operations [3]. Cutting fluids are customarily used to
control the temperature in the cutting zone. However, the
use of cutting fluids in machining processes has been
questioned, due to some adverse effects which they cause.
Prolonged contact of machine operators with cutting fluids
may cause skin and respiratory diseases. Improper disposal
of cutting fluids results in ground, water, and air pollution.
In addition, the costs related to cutting fluids are higher
than those related to labor and overhead. Thus, environmental and resource problems are forcing industry to
implement strategies to reduce the use of cutting fluids in
their production activities [4]. With this purpose, some
alternatives have been sought to minimize or even avoid the

use of cutting fluid in machining operations. Two of these
alternatives are dry machining and machining with the
minimum quantity of lubrication (MQL) [5, 6]. Among the
processes in which these techniques are being applied is
drilling.


432

Drilling is one of the most demanding machining
processes because a completely machined geometry and
surface are generated in one operation and usually
postmachining is impossible. The demands in regard to
diameter precision, straightness, and surface quality are
enormously high. Tools must meet the requirements for
diameter tolerances and shape–position tolerances [7]. In
dry drilling, tool failure is a significant factor affecting
productivity and manufacturing efficiency. Hence, one of
the main objectives of cutting research is the assessment of
tool wear and increasing tool life [8].
Beyond tool wear, the surface quality of the machined
components plays a key role. The ability of a material to
withstand severe conditions of stress, temperature, and
corrosion depends on the quality of the surface generated
during machining, which consequently determines the
longevity and reliability of products made of these
materials. The machined surface quality can be defined by
two measures: surface topography and subsurface integrity.
The surface topography can be measured using standard
surface roughness measurement equipment, whereas the

measurement of subsurface integrity is a complex task [9].
In a technology transfer from conventional machining
with abundant lubricant to dry machining or with MQL, the
thermal behavior tends to be more pronounced. Several
studies [10–14] present results which show the tendency
for a higher maintenance of the elevated temperatures in the
cutting zone when dry machining is applied. These studies
also indicate the tendency of lower temperatures using
abundant fluid and intermediate temperatures applying
MQL.
Previous works have shown good results for reducing or
removing the cutting fluids in drilling processes. Rahim and
Sasahara [12] conducted drilling experiments under different
cooling and lubrication conditions such as air blow,
vegetable and synthetic MQL, and flood. They used
coated (TiAlN) carbide drills in the drilling of titanium
alloy Ti-6Al-4 V. The researchers found that MQL gave
comparable performance with the flood condition. Bhowmick
et al. [10] also reported similar performance between MQL
and flooded drilling in the machining of cast magnesium
alloy AM60 using HSS drills. They used mineral oil in flood
condition, and two types of MQL fluids, distilled water and a
fatty acid-based. Heinemann et al. [15] demonstrated the
effect of MQL on the tool life of coated and uncoated HSS
twist drills. They performed deep-hole drilling experiments
in plain carbon steel and found that a low viscous MQL oil
with a high cooling capability gave rise to a notably
prolonged tool life. Bhowmick and Alpas [16] studied the
performance of diamond-like carbon (DLC)-coated HSS
drills under MQL condition in the machining of an Al–6%

Si (319 Al) alloy. Results were compared with drilling using
conventional flooded coolant. They applied two types of

Int J Adv Manuf Technol (2012) 58:431–441

DLCs (nonhydrogenated and hydrogenated) and distilled
water as the MQL agent. The MQL cutting using either type
of DLC coating reduced the drilling torque compared to dry
drilling to a level similar to the performance under the
flooded condition. Tasdelen et al. [3] made drilling experiments with MQL at different oil amounts, dry compressed
air, and emulsion. The holes with 33-mm depth and 19-mm
diameter were drilled with 155 m min−1 cutting speed and
0.11 mm rev−1 using indexable inserts. The tests showed that
MQL and compressed air usage have resulted in lower wear
both on the center and periphery insert compared to drilling
with emulsion.
It is well established that the main problems in drilling
when the cutting fluid supply is reduced or removed are the
higher maintenance of elevated temperatures in the cutting
zone and the difficulty for removing the chips from the hole
[6, 17–19]. This second aspect is especially critical for
drilling, since the cutting process is involved by the
material of the piece. When the chip flow is compromised,
it leads to packing and clogging of the chip and can cause
the collapse of the tool [20–23]. It is also known that these
problems are caused by the loss of the primary functions of
the cutting fluids, which are lubrication, cooling, and
transport of chips [6, 18, 19]. However, it is not clear what
are the changes in the interface tool/piece/chip due to the
loss of these functions and how they affect the tool wear

and the quality of the machined surface. Therefore, in view
of the complexity and extent of difficulties and different
conditions in this type of change process, this work presents
a study of drilling of AISI P20 steel with carbide tools, with
different conditions for the application of cutting fluid. The
main goal was to evaluate the effects on the process caused
by the reduction of the cutting fluid supply and its relation
with the tool wear and the surface integrity of the piece.

2 Experiments
2.1 Workpiece
The workpieces were prepared with AISI P20 steel and
were hardened by heat treatment to obtain a final hardness
between 36 and 38 HRc. This steel is frequently used in the
manufacture of molds and die cavities. The chemical
composition is given in Table 1.
The workpiece dimensions were 250×80×60 mm. The
distance between holes was 1.5 times the diameter of the
Table 1 Chemical composition of AISI P20 steel (% wt., ASTM)
C

Si

Mn

Cr

Mo

Ni

0.35–0.45 0.20–0.40 1.30–1.60 1.80–2.10 0.15–0.25 0.90–1.20


Int J Adv Manuf Technol (2012) 58:431–441

433

tool. In the dry tests, a drilling sequence was used with a
distance between holes equal to three times the diameter of
the tool using, however, two such cycles to complete the
workpiece. Thus, at the end of the second cycle, the same
distance between holes of 1.5 times the diameter of the tool
was obtained. This strategy was applied in the dry tests to
avoid thermal influences that could compromise the results
of the experiment.
2.2 Tools
The tools used in the experiments were coated carbide
drills, DIN 6537 K, provided by Walter AG Company. The
diameter of the tools is 8.5 mm and they are coated with
TiAlN. For the dry tests, some drill flutes were polished
with an abrasive cloth to obtain a smoother flute surface
and improve the chip flow. Figure 1 shows the standard tool
used in the experiments.
2.3 Equipment
The experiments were performed on an Okuma Ace Center
MB-46 VAE Vertical Machining Center, with maximum
rotation of 15,000 rpm and power of 18.5 kW. A Universal
stereoscope was used in wear analysis and measurements.
The same equipment was used for an optical analysis of the

texture of the machined surfaces. The surface roughness,
Ra, parameter, was measured using a Taylor Hobbson 3+
surface roughness tester. To analyze the microstructures and
to measure the depth of plastic deformations, a Nikon
Optical Microscope Epiphot 200 was used. Microhardness
tests were carried out with a Shimadzu HMV-2 microhardness tester to determine if there was any metallurgical
alteration in the subsurface region of the machined material.
2.4 Experimental procedures
The cutting parameters used in tests were a cutting speed of
50 m min and a feed of 0.1 mm. The hole depth used was
three times the diameter of the tool (25.5 mm). The tests
were carried out for three different quantities of cutting
fluids: (1) the application of fluid in abundance, (2) the
MQL, and (3) a total absence of fluid. For the emulsion, a
pressure of 3 bars was applied with a flow rate of 1,800 l
h−1. The oil used was Vasco 1000, in a concentration of
Fig. 1 Drill used in the
experiments

Type:
Grade:
Standard:
Diameter:
Coating:
N º of edges:

10%. In the MQL condition, the same pressure of 3 bars
was used with flow rate of 10 ml h−1. The MQL oil used
was VASCOMILL MMS SE 1. Both oils were provided by
Blaser Swisslube of Brazil.

The flow rate applied in MQL tests is the default value of
the machine used in the experiments. Some of the first works
dealing with MQL [24, 25] used to apply higher flow rates,
up to 300 ml h−1, but several studies showed that lower
values tend to present similar performance. Bhowmick et al.
[10] analyzed the average torque responses when MQL
fluids were supplied at the rates of 10, 20, and 30 ml h−1,
and they did not find significant difference. Braga et al. [26]
found similar results for tool wear using 10, 30, and 60 ml
h−1 of oil. Therefore, the flow rate of 10 ml h−1 was fixed in
this work.
Figure 2 shows the MQL system coupled to the
machining center and also details of the nozzle position
regarding the tool.
The quality surface analysis made in the holes was
carried out near the beginning of the hole and near the
bottom of the hole. The roughness measurements were
made at three equidistant points for each depth (initial and
final regions). Figure 3 shows the analysis depths and
illustrates the roughness measurement positions.
Three repetitions were made for each condition of fluid
application in order to get a satisfactory result. The end of
tool life criterion adopted was a maximum flank wear
(VBmáx.) of 0.2 mm or the occurrence of chipping. The tests
were interrupted after 1,200 holes, even if the tool did not
reach the end of life criterion.
For the application of fluid in abundance, the strategy of
continuous drilling was adopted. And for the MQL and dry
conditions, a pecking cycle was used, with an advance of
1.5 mm followed by retreat out of the hole. This procedure

was used to facilitate the expulsion of the chip from the
hole and to avoid crushing the chips and clogging the drill.

3 Results and discussion
3.1 Wear
The cutting parameters used in the tests were determined
through the analysis of cutting performance during preliminary experiments. In these tests, no cutting fluid was

carbide drill ALPHA 2
K30F
DIN 6537K
8.5 mm
TiAlN
2

measures
in mm

89
40

47
10

8.5


434

Int J Adv Manuf Technol (2012) 58:431–441

Fig. 2 MQL system (a) and
application position (b)

30°

a
Nozzle
position

14
mm

36
mm

MQL
system

b

60 mm

near the
beginning
Analyzed
depths

22 mm
2 mm

applied, increasing the process severity, to facilitate the
occurrence and observation of problems.
One of the main problems in dry drilling is the removal
of chips from the hole. This problem was observed in
continuous drilling using the cutting parameters vc =40 m
min−1, f=0.10 mm, when the maximum number of drilled
holes obtained was 129. In this condition, chip flow was
difficult and obstruction of the drill flute occurred, as seen
in Fig. 4a. Also, microchippings were observed in the
margins and the consequent chip removal from the flute
through the margin and land, as seen in Fig. 4b. The
combination of these problems compromised the chip flow,
causing the obstruction of the flutes and led to tool
breakage.
To improve chip flow, the drill flute of some tools was
polished with an abrasive cloth to obtain a smoother surface
and, in this way, to facilitate the expulsion of the chip.

1

2

3
near the
bottom

Fig. 3 Points of measurement

Roughness

measurement
points

MQL system: EcoBooster Model EB-3
Oil: Vascomill MMS SE 1
Work pressure: 3 bar
Flow rate: 10 ml/h

However, better results were not obtained and again
microchipping occurred in the margins, as seen in Fig. 5.
An attempt to solve the microchipping at the margins
was made, maintaining the same values of cutting speed
and feed, but adopting a pecking cycle, with an advance of
1.5 mm without retreat out of the hole. The intent with this
strategy was to improve the chip breakage and therefore
enhance the chip flow. But again, good results were not
obtained. The drill presented high adhesion of material in
the flute and in the margin, and a high packing factor
(clogging of the drill by the chips) was observed. Figure 6
illustrates this condition.
Therefore, a more detailed analysis of the chip formation
was necessary. For this purpose, chip samples were
prepared for metallographic analysis. Chips adhering to
the drill flutes were removed and mounted in order to
analyze their transverse section. This process started from
the chip region near the top of the drill, and the analyzed
depth increased successively by 1 mm after each analysis.
Between 3 and 6 mm from the top, a higher packing factor
was observed. This region is the same as that where the
microchipping of the margins occurred. Figure 7a shows

schematically the mechanics of cutting and shear zones, for
a better understanding of the results presented in Fig. 7b,
which shows the transverse section of the chip on the depth
of 3 mm below the top of the drill.
Figure 7b shows the disordered packing of the chip and
microwelding points as well as the different regions
(smoothed and sheared) resulting from chip formation.
After severe machining, the material removed can present
several metallurgical alterations such as a high plastic
deformation, a hardness increase, or a white layer formation. The sheared region resulted from the primary shear


Int J Adv Manuf Technol (2012) 58:431–441
Fig. 4 Flute obstruction (a) and
microchipping in the margin (b)

435

a

b

5 ho le s

2 mm

2 mm

zone on cutting and, therefore, is submitted to severe strain
hardening. The smooth region corresponds to the secondary

shear zone, and the material is submitted to high compression and attrition against the tool face. Under severe
conditions of pressure and temperature, this region can
develop the so-called white layer, characterized by high
plastic deformation and high hardness. Thus, smooth and
sheared regions present metallurgical alterations that tend to
increase the resultant hardness [28]. And this increase in
hardness of the chip makes it more difficult to flow through
the flutes on its way to removal from the hole.
The smooth and sheared regions, by having a high
mechanical strength and due to the forces generated by
compression and mechanical wear, tend to increase the heat
generation. Thus, the three necessary conditions for the
formation of a white layer on the chip are present:
compression, friction, and high temperature [29]. The
friction of the regions of high hardness with the hole wall
creates difficulty for disposing of the chip, resulting in a
volume increase, which compromises the chip flow out of
the hole. The chip accumulation causes the obstruction of
the flutes and the consequent tool breakage.
However, although this analysis has shown the occurrence of high compression and friction loads on the chip, it
did not explain the cause of these problems. Then, a
complementary analysis was made with a broken drill
inside a hole. An external cylinder of material involving the
hole with the tool and chip was removed from the
workpiece. This cylinder was prepared for optical and
Fig. 5 Microchipping in the
margin, external (a) and internal
(b) view

3 0 hole s

a

metallographic analyses with a special cold cure resin that
enabled transverse cuts of the cylinder in 2-mm-thick
layers, maintaining the position of the tool and the chip
inside the hole (see below). The aim of this analysis was to
evaluate the chip formation and the interface tool/chip/piece
and to investigate the cause of the elevated packing factor
of the chip. Figure 8 illustrates the procedure and shows the
top surface of the layer cut from approximately 1.5 mm
below the top of the tool.
The layers were carefully analyzed and an unexpected
cutting behavior was observed. Figure 8 shows that the chip
has been cut by the margin, but these tool elements are not
designed for cutting. The hypothesis raised for this case
was that the wear of the drill corners would lead to a
reduction of the tool diameter between the corners, and thus
the diameter between the margins would be larger and the
cutting of the material would be made by the margins. This
difference between the diameters makes the margins the
primary cutting surfaces of the drill and implies a change in
the shear planes regarding the feed axis, transferring the
primary shear plane, shown in Fig. 7a, to the margin. This
change makes the chip flow toward the hole wall, instead of
following the helical path along the flute. This results in
intense friction with the hole wall and elevated compression
loads on the chip, giving rise to the observed high packing
factor of the chip and the microchippings in the margins.
These results explain the lower results for tool life in dry

continuous drilling experiments. Based on these analyses, it
was decided to change the drilling strategy in order to
External view

b

Internal view
margin/flute

85 holes
Adhesion

Substrate
1 mm

0.75 mm


436

Int J Adv Manuf Technol (2012) 58:431–441

Fig. 6 High packing factor (a)
and material adhesion (b)

a

b

3 holes

2 mm

facilitate the expulsion of the chips from the hole.
Therefore, it was adopted as the pecking cycle with an
advance of 1.5 mm followed by a retreat out of the hole.
The retreat improved the chip expulsion from the hole, and
in this way, the microchippings at the margins no longer
occurred. However, this strategy increased the drilling time
by about 150%. To compensate part of this loss of
productivity, tests were made with the cutting speed of
50 m min−1, which showed similar results to those obtained
with 40 m min−1. Thus, for the main experiment, the
cutting speed of 50 m min−1 was adopted, and the pecking
cycle for MQL and dry conditions was used. For processing
with the emulsion, the tests were performed by continuous
drilling in order to attain lower cutting times. With these
definitions, the main experiment was carried out, and the
results are presented below. Figure 9 presents the wear
results from the experiments with vc =50 m min−1.
Drilling with the emulsion gave the worst result and the
tools made 933 holes on average. For MQL tests, all three
tools made 1,200 holes, but only one reached the end of
tool life. For the other two, the test was interrupted by the
criterion of 1,200 holes drilled. The dry drilling experiments presented the best results, because all tests were
interrupted by the 1,200 holes criterion.

a

0.75 mm

The worst results for the emulsion tests can be explained
by the cooling of the machined material. The presence of
the cutting fluid removes the positive effect of the heat in
the cutting zone, which facilitates the material shear (lower
resistance to cutting). That way, the cooled material
presents greater strength, increasing the mechanical loads
on the tool and, consequently, the tool wear. In the emulsion
tests, also a high adhesion of material on the flank of the
tools was observed, which leads to microchipping and tool
failure. Figure 10 shows the material adhesion and the
consequent microchipping.
The tools machined under MQL conditions also presented material adhesion on the flanks, but in lesser
quantities than observed for the emulsion condition. In
dry tests, it was observed that although material adhesion
had occurred on tool flanks, when this material detached
from the flank, it did not cause significant microchipping as
was observed in the emulsion and MQL experiments.
3.2 Surface quality
The quality of machined components is currently of high
interest, for the market demands mechanical components of
increasingly high performance, not only from the stand-

b

Shear angle

Shear plane
Primary shear zone

Smooth
region

Chip
Piece

Sheared
region

Rake
angle

Micro-welding
Tool
Secondary shear zone

Fig. 7 Shear planes [27] (a) and the transverse section of the chip distant approximately 3 mm from chip region near the top of the drill (b)


Int J Adv Manuf Technol (2012) 58:431–441

437

2 mm

Tool
Chip

Margin

Nital 2%

1.5 mm

Piece

Fig. 8 Interface piece/tool/chip analysis

point of functionality but also from that of safety.
Components are produced through operations involving
the removal of material display surface irregularities
resulting not only from the action of the tool itself but
also from other factors that contribute to their superficial
texture such as cutting speed, tool wear, feed, tool
materials, tool geometry, etc. This texture can exert a
decisive influence on the application and performance of
the machined component [30, 31].
With the aim of facilitating the comparisons, in the
surface quality analysis made for this experiment, one
representative tool was selected and tested for each
condition of fluid application that made 1,200 holes without
attaining the end of tool life criterion. Figure 11 shows the
values of roughness, Ra, measured near the beginning and
the bottom of the holes, for each drilling condition.
Near the beginning of the holes, the dry tests gave the
highest roughness values, while the lowest values were
measured in holes machined with MQL. The elimination of
cutting fluid tends to worsen the quality of the surface due
to the larger friction forces and the increased detachment of
material particle adhesions that are released from the tool

[32]. For the emulsion condition, the fluid reaches the
initial region of the hole, providing good cooling and
lubrication, and for MQL condition, with the employment
of the pecking cycle, the edge receives a microlubrication

a

b
vc = 50 m/min f = 0.1 mm

1400
1200

Number of holes

Fig. 9 Picture of the drilling
process (a) and wear results
expressed in terms of the
number of holes drilled (b)

after each retreat, which reduces the attrition and produces
lower roughness results. However, after about 1,000 holes,
the results of all lubricant conditions tend to converge
toward similar values. The friction caused by dry drilling is
reduced while producing the holes due to the adjustments
of the cutting edge, decreasing the roughness in the
machined surface.
The MQL conditions also resulted in the lowest
roughness values in the analysis made near the bottom of
the holes. However, for this analysis condition, the

emulsion tests presented a tendency toward larger roughness values, because the emulsion condition keeps the
original cutting edge geometry for a greater time, causing
the increased roughness. But after approximately 1,000
holes, the same tendency to similar values for all conditions
is observed, as it was in the previous analysis. Texture and
microhardness analyses made near the bottom of the holes
showed, especially for dry condition, the occurrence of
microwelding of the chip on the surface, which can be
caused by elevated temperatures during machining resultant
from the worn cutting edge. A worn cutting edge has its
geometry changed, which reduces its cutting ability,
hindering the shear of the material. With that, due to
friction and the high temperatures generated in the process,
parts of the removed material are welded onto the surface,
providing a smooth aspect, which reduces the values of

1000
800
600

Continuous
drilling

Pecking
cycle

Pecking
cycle

Emulsion

MQL

Dry

400
200
0


438

Int J Adv Manuf Technol (2012) 58:431–441

Fig. 10 Flank images with
material adhesion (a) and
microchipping (b)

Emulsion
Drill after 1000 holes

Emulsion
Drill after 800 holes

a

b

1 mm

roughness [33]. Figure 12 shows the surface texture of the
first and last holes for the different fluid application
conditions.
For drilling with the emulsion, the marks of the passage
of the cutting edge can be seen, and in the first hole, there
are deeper grooves which result in greater roughness values
than in the last hole. The holes drilled with MQL presented
a more homogeneous texture along the holes, especially in
the last hole. For the MQL condition, the microlubrication
associated with the facilitated material shear due to the
temperature increase in the cutting zone reduces the cutting
forces and allows the formation of a smooth surface. The
holes obtained by dry drilling presented the worst visual
aspect, with evidence of material adhesion along the holes
and a smoothed surface in the final region of the holes,
probably due to the occurrence of microwelding.
To complement this analysis, Fig. 13 illustrates the
surface texture of the region near the bottom of the 1,200th
hole with a larger zoom and also shows the average value
of the microhardness measured on the surface of the hole
for each condition. The measured values are approximately
twice as large as the bulk material hardness, 390 HV on
average. These results corroborate the hypothesis of the
occurrence of microwelding, since the welded chip tends to
be submitted to high thermal and mechanical loads, which
can lead to the microhardening of the chip.

1 mm

As mentioned previously, none of the tools analyzed for

surface quality attained the end of life criterion, and the
tests were interrupted after 1,200 holes drilled by each tool.
However, wear analysis showed the maximum flank wear
to be 0.15 mm for the tool used in the emulsion condition,
0.08 mm for the tool used under dry conditions, and
0.06 mm for the tool used in the MQL condition.
Considering that the tool has a cutting edge radius of
0.05 mm, it can be stated that these wear values have a
significant influence on the surface changes.
Beyond the surface region of the holes, the metallurgical
alterations in the surface integrity were also studied.
Surface quality influences characteristics such as fatigue
strength, wear rate, corrosion resistance, etc. The fatigue
life of a machined part depends strongly on its surface
condition. It has long been recognized that fatigue cracks
generally initiate from free surfaces. This is due to the fact
that surface layers experience the highest load and are
exposed to environmental effects. Crack initiation and
propagation, in most cases, can be attributed to surface
defects produced by machining. The surface of a part has
two important aspects that must be defined and controlled.
The first aspect involves the geometric irregularities on the
surface, while the second aspect involves metallurgical
alterations of the surface and the subsurface layer. The latter
has been termed surface integrity [34].

3.0

Dry

2.5

2.5

2.0

2.0

Region near the
beginning of the hole

1.5

Emulsion

1.5

Dry

1.0

Emulsion

Number of holes

1

00
11
00

12
00

10

0
70
0
80
0
90
0

60

0

0

0
50

40

30

0

0
20

10

1

0.0

Number of holes

Fig. 11 Graphs of roughness Ra values measured near the beginning and near the bottom of the holes

0.5

MQL

0.0

00
11
00
12
00

Region near the
bottom of the hole

MQL

0
20

0
30
0
40
0
50
0
60
0
70
0
80
0
90
0

0.5

10

1.0

10

Roughness Ra (µm)

3.0


Int J Adv Manuf Technol (2012) 58:431–441

439

Emulsion
Hole 1200
Hole 1
Ra = 1.07 µm

3 mm

Ra = 1.49 µm

Dry

MQL

Ra = 0.60 µm

Hole 1

Hole 1200

Hole 1

Hole 1200

R a = 0.89 µm

Ra = 0.56 µm

Ra = 0.78 µm

R a = 0.65 µm

3 mm

3 mm

Ra = 0.43 µm

R a = 1.03 µm

3 mm

Ra = 0.69 µm

3 mm

Ra = 1.22 µm

3 mm

R a = 0.57 µm

Fig. 12 Surface texture of first and last holes for the different fluid application conditions

Surface integrity analysis has great importance for
surface quality characterization, due to its direct relation
with the performance of the machined component. This
characterization of the integrity can be performed by the

evaluation of the alterations of the structure under the
surface, as measurements of plastic deformations, microhardness, among others. Plastic deformations consist in the
deformation and change of orientation of the grains near the
surface of the material after the cutting. The measured
values correspond to the vertical distance from the surface
to the point in the microstructure without visible alterations.
Figure 14 presents the values of the plastic deformation
measured near the beginning and the bottom of the first and
Fig. 13 Surface texture in the
region near the bottom of the
last holes

last machined holes. Each value plotted in the graphs is the
average value of the five maximum plastic deformations
found in the analyzed region. The figure also shows
metallographic images of the last hole surface for the
different conditions tested.
For both analyzed regions (near the beginning and near
the bottom), the measured deformation values were larger
with the increase in the number of drilled holes due to the
changed cutting edge geometry; this tends to increase the
temperature in the cutting zone and, in turn, leads to the
occurrence of higher deformations. It can also be observed
that the measurements made near the bottom tend to be
greater than those made near the beginning of the holes.

Region near the bottom of the 1200th hole

775.4 HV

934.8 HV

846.2 HV

0.1 mm

0.1 mm

0.1 mm

Dry

MQL

Emulsion

Ra = 0.69 µm

Ra = 0.43 µm
2 mm

Ra = 0.57 µm
2 mm

2 mm


440

Int J Adv Manuf Technol (2012) 58:431–441

1200th
hole

Average plastic
deformation [µm]

50

emulsion

dry

MQL

emulsion

Region near the beginning of the hole

40
30

dry

Region near the bottom of the hole

Dry

Dry

Emulsion

Emulsion

MQL

MQL

20
10

MQL

0

1st hole

1200th hole

1st hole

1200th hole

Fig. 14 Average plastic deformation measured near the beginning and near the bottom of the holes

Fig. 15 Surface microhardness
versus the distance from
machined surface

Micro-hardness [HV]

This can be related to the greater difficulty of chip removal
from the hole in the final region, increasing the contact
between the chip and the piece, which causes the increase of
heat generation and consequently higher deformation values.
Also, for the beginning and final regions, the dry tests
resulted in greater plastic deformation, due to the process
severity, because, as stated by the literature, this condition is
characterized by higher maintenance of elevated temperatures in the cutting zone [6, 17–19]. Near the beginning of
the holes, emulsion and MQL results were very similar,
while near the bottom of the holes, the emulsion deformation results were larger. This difference is explained by the
cutting strategy adopted for each condition. In the initial
region, the fluid can reach the cutting zone for both
conditions. However, in the final region, the continuous
drilling employed in emulsion test impeded fluid access to
the cutting zone, while in the MQL test, with the
employment of the pecking cycle strategy, the edge receives
a microlubrication after each retreat, which reduces the
temperatures generated in cutting and results in lower
deformations values.
To complement the subsurface analysis, microhardness
measurements were performed. Because of equipment
limitations, the first possible distance for measurement is
0.02 mm from the machined surface. As seen in Fig. 15,
which presents the microhardness measurements made near
the beginning and the bottom of the last machined hole for
the different fluid application conditions, the measured
460

results presented a normal dispersion around the bulk
material hardness, and no significant trend was observed.

4 Conclusions
The preliminary tests for the dry condition with continuous
drilling presented great difficulty for chip flow through the
drill flutes, generating a high chip-packing factor. These
conditions led to obstruction of the flutes and the
consequent tool breakage. After an analysis of the chip
formation, it was concluded that the cutting was being done
by the margins, instead of the principal edges, which
changed the shear planes regarding the feed axis, and
compromised the chip flow along the flutes. To resolve this
problem, the pecking cycle was adopted, with a periodic
retreat out of the hole. This strategy improved the chip flow
and stopped the tool breakages. Thus, the main experiment
was performed with continuous drilling for emulsion tests
and with the pecking cycle for MQL and dry tests.
The surface quality analysis showed that, near the
beginning and the bottom of the holes, the dry drilling
condition generated greater values of the roughness on the
machined surface due to the higher friction on the interface
tool/chip/workpiece caused by the absence of the coolant
and lubricant functions performed by the cutting fluid. The
MQL application condition resulted in the lowest roughness
values in both analyzed regions of the holes. Near the

Region near the beginning
of the 1200th hole

Region near the bottom
of the 1200th hole

Emulsion
MQL

420

460

420

Dry
380

380

Bulk
material
hardness

340
0

20

40

60

80

100

Distance from the surface [µm]

120

0

20

40

60

80

100

Distance from the surface [µm]

340
120


Int J Adv Manuf Technol (2012) 58:431–441

beginning of the holes, differently from what occurred in the
region near the bottom, the emulsion tests presented a
tendency toward higher values of roughness, because emulsion keeps the original cutting edge geometry for longer times
and also cools the machined material, keeping its shear

strength high and causing well-defined grooves due to the
passage of the tool and the consequent greater roughness.
The microhardness measurements on the surface of the last
holes, in the region near the bottom, resulted in values
approximately twice as large as the bulk material hardness,
which corroborates the hypothesis of the occurrence of
microwelding, because applying high thermal and mechanical
loads to the welded chip causes microhardening of the chip.
Acknowledgments The authors would like to thank Walter A.G.
Company, Blaser Swisslube do Brasil, and Okuma Latino Americana
Comércio. They also wish to thank Prof. Frank P. Missell for useful
discussions. This work was supported by the Brazilian agency CNPq.

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Int J Adv Manuf Technol (2012) 58:443–463
DOI 10.1007/s00170-011-3431-2

ORIGINAL ARTICLE

Virtual workpiece: workpiece representation for material
removal process

Seok Won Lee & Andreas Nestler

Received: 8 September 2010 / Accepted: 20 February 2011 / Published online: 5 October 2011
# Springer-Verlag London Limited 2011

Abstract In this paper, an efficient methodology to
generate a virtual workpiece (VWP) is presented. VWP is
a workpiece in a virtual environment in which the
geometric, kinematic, and thermo-mechanical effects of
the process and resources can be reflected. VWP encompasses not only the macro-information corresponding to the
shape of the “virtually” machined intermediate workpiece,
but also the micro-information, such as the surface
roughness, scallop heights, chatter mark, etc. To represent
VWP, swept volume (SV) of geometrically defined cutters
is generated first by envelope profiles which are calculated
by the intersection of the Tool map with the Contact map of
the tool moving direction. Then SV is tessellated to conduct
elementary 1D Boolean subtraction of SVs from the IPW.
The Boolean subtraction is realized by means of an efficient
ray-triangle intersection test using Barycentric coordinates.
Finally, VWP is reconstructed as a triangular mesh (STL,
stereolithography data format) from the orthogonal tripledexel model (TDM) which predicts machined surface
quality, such as surface roughness, gouging and sharp
edges and is reused for further operations, e.g., tool path
generation, simulation and geometric metrology, etc. To
demonstrate the validity of VWP modeling, several material
removal processes, e.g., milling and micro-EDM operations, have been tested and the proposed approach has been
proven to be applicable to enhance the quality of NC
simulation and verification.

S. W. Lee (*) : A. Nestler
Institute of Forming and Cutting Manufacturing Technology,
Dresden University of Technology,
01062 Dresden, Germany
e-mail:

Keywords Virtual workpiece . Material removal .
Machining simulation . Volume updating . Five-axis
machining

1 Introduction
A numerical control (NC) milling removes metal or other
material out of a stock by moving a cutting tool
intermediate space. Because the NC programs, which
contain the control commands of the machine, are usually
not always error-free and time to market is getting shorter
in modern production cycles, verification of milling
operations plays a crucial role in manufacturing processes
[1]. Since the cutting simulation is the material removal
process with a geometrically defined cutting tool, which is
equivalent to the Boolean subtraction of cutter volumes
from raw stock. It is essential to continually subtract the
exact swept volume(SV) of tool from the raw stock in the
virtual environment in order to predict final shape as
realistically as possible. Performance of the Boolean
subtraction depends on execution time, accuracy, and the
storage (memory) needs. There have been many attempts to
verify material removal processes such as milling prior to
real machining and, as a result, many representation
schemes of the workpiece model exist. For instance, there

are workpiece models based on CSG, Z-map, dexel, voxel,
NURBS and vector representation, etc. Hereafter, the
workpiece representation methods applied for the purpose
of the milling simulation are presented.
Z-map representation [2, 3] is the most prevailing
method because of its simplicity and robustness, but it is
impossible to simulate the machining of the undercut
geometry, such as overhang shape. To overcome the


444

shortage of Z-map representation, Jerard et al. [4] uses the
discrete vector model (DVM; or lawnmower method) that
is represented by the surface normal vectors at discrete
surface points. In this model, if the cutting tool moves near
to the part surface, the vector is shortened, which
corresponds to the material cutting. DVM is efficient for
the finishing operation where the finishing allowance is
removed from the design surface so that error assessment is
possible. However, it is inadequate for roughing operations
where an intermediate workpiece is to be calculated. Van
Hook [5] extends the Z-map representation to accomplish a
five-axis milling simulation. Each dexel or depth element is
defined as a rectangular solid extending along the Z-axis
and consists of a near and far Z depth, a color, and a pointer
to the next dexel. A dexel has the Z value of the farthest
surface as well as that of the nearest one and is linked with
the neighboring one. This difference makes it possible to
visualize the undercutting during five-axis machining in

image space. The dexel algorithm can be implemented in
read, write, comparison, and conditional branches without
geometrical complex computations. Hence, it is fast and
robust so that it has been widely adopted in a lot of
simulation software [6, 7]. Huang and Oliver [8] develop an
NC milling verification system by extending the aforementioned dexel approach, which enables the interactive
visualization of milling simulation and modification of
five-axis milling tool paths by means of an assessment of
dimension errors.
Voxel representation [9, 10] has advantages over other
representations in that the Boolean set operations are
conducted at the level of primitive volumetric element or
voxel, and so it reduces the computational complexity of
regular Boolean set operations. Karunakaran and Shringi
[11] implement a solid model-based volumetric NC
simulation system using octree representation, which is an
adaptive version of the voxel representation and subdivides
the space into eight parts recursively to simulate cutting
process and optimize the cutting parameters.
Benouamer and Michelucci [12] advocate a multidexel model (MDM), which consists of three conventional
dexel models orthogonal to each other, and the marching
cubes (MC) algorithm to implement an NC milling
simulation module [13]. Each dexel contains the entering
and exit hit (real value), material IN/OUT information
(binary value) and, if any, the material property. The
cutting tool is sampled along each axis for each interpolated time and the hit points are updated in MDM. The
method is simple and easy to implement, but has many
considerable limitations. First, several numerical and
topological inconsistencies caused by the multiple usage
of the single-dexel model (SDM) are to be clearly solved

under a certain presumption. Additionally, there are
aliasing and under-sampling problems: if an initial

Int J Adv Manuf Technol (2012) 58:443–463

sampling size is too large it may lead to overlooking a
small object or featured areas such as sharp edges and
sharp corners. Therefore, determining an optimal sampling
size of an object as well as the huge storage management
required as compared to a SDM would be challenging
problems for MDM [14].
Gläser and Gröller [15] propose a robust algorithm to
update workpiece volume in the framework of polygonal
boundary representation at three-axis machining in order to
design optimal tool shapes and paths. Granados et al. [16]
implement a complete, exact and correct Boolean operation
via Nef polyhedra which is the closure of half-spaces, e.g.,
ax þ by þ cz < d. Thereby the exact arithmetic is applied
to overcome an incorrect rounding problem of the floatingpoint arithmetic [17]. The algorithm is implemented in
CGAL [18] and the resultant geometry after the Boolean
operation is precise and robust. However, the calculation
time is not amenable for milling application even though
the algorithm is optimized later in Ref. [19].
By applying a graphics processing unit (GPU) Bohez et
al. [20] developed sweep plane algorithm using the stencil
buffer to simulate the five-axis milling operation. In
addition to the color and depth buffers, the stencil buffer
is an extra buffer, which is usually used to restrict the
drawing area to certain portions of the screen [21]. Using it
they reduce 3D Boolean subtraction into a laminated 2D

plane Boolean operation. Inui and Kakio [22] exploit the
hidden surface removal technique, which corresponds to the
Z-map method, of GPU to perform the Boolean subtraction
of the workpiece. The algorithm eliminates surfaces which
are positioned at a farther distance and thus is invisible
from the viewpoint and saves color and depth value (or Z
value) of the nearest surface pixelwise. Saito and Takahashi
[23] introduce a simple and effective geometric buffer (Gbuffer) method, which is integrated in one methodology, to
simulate three-axis NC machining [24].
From the investigation of workpiece representation models, it could be concluded that the simulation of five-axis
machining is still a challenging problem despite much
achievement by many researchers. This is in last decades
because (1) the exact SV generation is a critical hurdle to
update the workpiece during five-axis machining due to the
geometrical and computational complexity. The cutting
operation of SVs from a workpiece model is still a bottleneck
in the NC simulation [25]; (2) representation methods of SV
and workpiece model are diverse so that the optimal
combination of them cannot be determined with ease; (3)
the process of updating the workpiece is composed of three
calculating steps, i.e., SV generation, Boolean operation, and
visualization. The identification and optimization of computational bound for each step is still a challenging theme so
that the optimal distribution of the computation burden
among them is not clearly reported [26].


Int J Adv Manuf Technol (2012) 58:443–463

445

and is worth considering together with the development of
data models [28].

2 Virtual workpiece
A practical NC simulation should guarantee the prediction
of potential errors prior to real cutting via a workpiece
model. However, the robustness of the Boolean operation
and its execution time has a reciprocal relationship. Thus
the efficient implementation of the Boolean operation and
the workpiece representation is a crux in NC verification. In
this section, the virtual workpiece, which is the main focus
through the work, is introduced based on the related work
addressed in Section 1.
2.1 Definition of virtual workpiece
The virtual workpiece (VWP) is the workpiece in a
simulation environment in which the geometric, kinematic
and thermo-mechanical effects of the process and resources
are reflected. VWP is influenced by the machine-cutterworkpiece system, CAM/CNC systems, machining operations, etc. (see Fig. 1). It contains not only the macroinformation corresponding to the shape of the “virtually”
machined intermediate workpiece but also the microinformation such as the surface roughness, scallop heights,
chatter mark, etc. VWP stores such surface information at
any step of simulation so that the analysis at every stage of
simulation is possible. The accuracy of VWP depends on
the geometric accuracy of SV of tools, mechanistic force
model of the cutting process and dynamics of the machine
tool in turn. Therefore, VWP is the core in which
comprehensive characteristics of the physical workpiece
are collected.
The goal “First part correct and fast” of virtual
machining could be reached by determining the extent of
how well the VWP could integrate the real machining

processes in the virtual manufacturing environment where
relevant databases, e.g., product, process, and resource are
tightly interlocked. VWP comprises the interrelationship
among the process parameters, material characteristics of
the workpiece and cutter, the type of interpolation, the
kinematic and kinetic parameters of machine tools, etc. [27]
Machining Operation Model

Cutting Simulation

Modeling of machining
operations with geometrically
defined cutter for:
- conventional and
- HSC machining

Interaction of cutter,
workpiece, clamping devices and
updating of in-process workpiece
ce
geometry

CAM

Virtual
Workpiece

Machine kinematics
- Number of axes: 3, 3+2, 5
Machining strategy

- tool path generation depending on
cutter types

Virtual Machine Tool
Thermo-dynamical model
of machine tool
for identifying influence factors
at machining

Fig. 1 Influencing area of virtual workpiece

2.2 Application fields of virtual workpiece
NC machine tools enable accomplishing machining
operations, in addition to achieving high accuracy and
productivity of machining parts. However, conventional
NC machine tools need an NC program, with which it is
difficult to change cutting parameters (e.g., depth of cut,
etc.) at machine level, to move axes. Moreover, the
cutting tools are to be pre-determined to generate part
programs properly before machining. Therefore, the NC
program is specified for the target machine tool and
cutting tools. The cause of this is a loss of flexibility of
machining operations.
The next generation NC controller is capable of adapting
and re-planning the machining schedules autonomously
depending on machining conditions and available resources
in the machine tool in comparison to the legacy controllers.
Adaptation and flexibility of machining and abnormal
condition avoidance can be achieved through a real-time
machining simulation. The VWP can play a significant role

in the next generation CAM/CNC systems, which enable
VWP to leverage synergic effects (flexible and intelligent
application fields (AF) of VWP) as follows:
1. Authentic five-axis simulation: the cutting simulation
becomes an indispensible tool in modern CAM systems
to cope with the mass customization which is characterized as adaptation and reconfiguration. While the SV
of three axis-machining is calculated in a simple
manner the SV of the five-axis machining is generated
very limitedly because the capabability of prevailing
geometric modeling kernels such as Parasolild [29],
ACIS [30], OpenCascade [31], etc. Even machining
simulation companies, such as MachineWorks (UK)
[32], Vericut (USA) [33], and ModuleWorks (Germany)
[34] are using the approximation of SV for five-axis
machining to increase the simulation speed at the
expense of the accuracy of simulation results. Three
main aspects, however—the accuracy, the calculation
speed, and the memory demand—need to be thoroughly considered in order to implement a realistic and
predictable machining simulation. Therefore accurate SV
calculation and the machining simulation based on SV are
the key features used to meet these requirements.
2. Intra-operability in machine tools: the interrupted
machining due to abnormal conditions (e.g., tool
breakage or tool wear) could be resumed if the VWP
is calculated and the tool path is regenerated based on
the VWP with available resources e.g., cutting tools in
the tool magazine.


446

3. Inter-operability between machine tools: this is a
similar case as with AF2, but the level of usage of
VWP is different. Owing to modification of a schedule,
or priority of orders, the current milling operation of
machine tool A should be halted and another order
should to be set up on the machine tool A. The halted
operation could be resumed on another machine tool B
which is available at that moment. The VWP on the
machine A is updated up to the halted time and used as
an input geometry, i.e., a raw stock for a machine B.
4. Geometry-based process monitoring: the commercial
milling process monitoring systems [48] show a
strength during mass production but a lack of valid
monitoring system of lot size 1. This results from the
fact that the old G & M codes [35] contain no
geometric information so that the contour is not known
a priori. On the other hand, the smart machine [49]
enables the geometry-based monitoring, which keeps
track of the VWP and distinguishes the areas where the
cutter begins to engage or exit the material.
5. Feed rate/tool speed optimization: VWP contributes to
the optimization of the cutting tool’s feed rate. The
actual volume of material removal eliminated from the
workpiece can be accurately calculated based on VWP.
This volume is the crux for the prediction of the cutting
forces and optimization of the feed rate by means of
mechanistic process models [36, 50].
6. Advanced tool path generation: today, the tool path
generation of modern CAM systems is based on the

solid/surface models which are constructed by CAD
systems. As usual, the security plane of the tool is
defined with constant distance from the stock material
surface so that the tool movement is not actually
economized. There is considerable room for rationalizing the air paths by using VWP. That is, the cutting tool
could cross over the area that has already been cut in
order to save cutting time. Moreover, tool path could be
calculated so that the engagement angle of the tool
could be kept constant by considering the updated
VWP.
7. Collision prevention: a data model for the machine tool
and the VWP enables the prevention of collisions
between the machine components and the spindle, tool
holder and tool itself in the controller. Additionally, the
collision check between VWP and other peripheries,
e.g., jig and fixture, etc., is especially useful for tool
path generation because the machine crash leading
to heavy personal and technical costs is preventable.
In the following sections, a methodology to generate
VWP during five-axis machining is explored. We
explain a cutting simulation kernel for five-axis machining which takes SV generated by the proposed approach

Int J Adv Manuf Technol (2012) 58:443–463

as input and then provide the practical implementation
of VWP which is necessary to accomplish each AF
(AF1-7).
This paper presents an efficient methodology to
perform a Boolean operation of triangulated SV in the
framework of triple-dexel model (TDM), to keep track of

VWP and to save VWP as polygon model for analysis of
the geometry and further usage. Main contributions in
this work are:
&
&
&
&
&
&

SV generation of conical tools undergoing five-axis
motion via Gauss map
Efficient methodology and algorithms to update the
in-process workpiece during simultaneous five-axis
machining using five-axis SV
The sampling technique of triangulated SV in an
efficient manner
Surface reconstruction from sampled data by recognizing feature sensitive geometries, e.g., sharp edges and
corners
Practical implementation of cutting simulation
Comprehensive performance analysis including detecting
bottleneck among modules through various examples

This paper is organized as follows: Section 3 provides
the SV representation method of prevailing cutters undergoing simultaneous five-axis movement via contact map
(C-Map) and tool map (T-Map). Section 4 presents VWP
model. Then the algorithms of the cutting operation in
the framework of the VWP model are elucidated in
Section 5. Section 6 discusses the surface reconstruction
from the sampled data. In Section 7, implementation

issues and performance analysis of the proposed approach
are addressed with various examples, and finally, in
Section 8 this paper is concluded.

3 Tool swept volume model
Lee and Nestler [37] have presented the method for
modeling SV of cylindrical tools that undergoes axisvarying motion via the Gauss map. This section explains
the basic idea and extends it to SV modeling of conical
tools undergoing simultaneous five-axis motion.
3.1 Contact map
Given an arbitrary vector τ ðkt k 6¼ 0Þ in the space, a great
circle is achieved by intersecting the unit sphere with the
plane orthogonal to τ passing the center of the sphere. This
great circle is called the C-Map of the generating vector τ
(Fig. 2). The C-Map CM of τ is defined as follows:
CM ¼ fnjn Á t ¼ 0; knk ¼ 1; kt k 6¼ 0g

ð1Þ