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ScienceDirect
Available online at www.sciencedirect.com
Available online at www.sciencedirect.com
ScienceDirect
Procedia CIRP 00 (2017) 000–000
www.elsevier.com/locate/procedia
2212-8271 © 2017 The Authors. Published by Elsevier B.V.
Peer-review under responsibility of the scientific committee of the 28th C IRP Design Conference 2018.
28th CIRP Design Conference, May 2018, Nantes, France
A new methodology to analyze the functional and physical architecture of
existing products for an assembly oriented product family identification
Paul Stief *, Jean-Yves Dantan, Alain Etienne, Ali Siadat
École Nationale Supérieure d’Arts et Métiers, Arts et Métiers ParisTech, LCFC EA 4495, 4 Rue Augustin Fresnel, Metz 57078, France
* Corresponding author. Tel.: +33 3 87 37 54 30; E-mail address: paul.stief@ensam.eu
Abstract
In today’s business environment, the trend towards more product variety and customization is unbroken. Due to this development, the need of
agile and reconfigurable production systems emerged to cope with various products and product families. To design and optimize production
systems as well as to choose the optimal product matches, product analysis methods are needed. Indeed, most of the known methods aim to
analyze a product or one product family on the physical level. Different product families, however, may differ largely in terms of the number and
nature of components. This fact impedes an efficient comparison and choice of appropriate product family combinations for the production
system. A new methodology is proposed to analyze existing products in view of their functional and physical architecture. The aim is to cluster
these products in new assembly oriented product families for the optimization of existing assembly lines and the creation of future reconfigurable
assembly systems. Based on Datum Flow Chain, the physical structure of the products is analyzed. Functional subassemblies are identified, and
a functional analysis is performed. Moreover, a hybrid functional and physical architecture graph (HyFPAG) is the output which depicts the
similarity between product families by providing design support to both, production system planners and product designers. An illustrative
example of a nail-clipper is used to explain the proposed methodology. An industrial case study on two product families of steering columns of
thyssenkrupp Presta France is then carried out to give a first industrial evaluation of the proposed approach.
© 2017 The Authors. Published by Elsevier B.V.
Peer-review under responsibility of the scientific committee of the 28th CIRP Design Conference 2018.
Keywords: Assembly; Design method; Family identification
1. Introduction
Due to the fast development in the domain of
communication and an ongoing trend of digitization and
digitalization, manufacturing enterprises are facing important
challenges in today’s market environments: a continuing
tendency towards reduction of product development times and
shortened product lifecycles. In addition, there is an increasing
demand of customization, being at the same time in a global
competition with competitors all over the world. This trend,
which is inducing the development from macro to micro
markets, results in diminished lot sizes due to augmenting
product varieties (high-volume to low-volume production) [1].
To cope with this augmenting variety as well as to be able to
identify possible optimization potentials in the existing
production system, it is important to have a precise knowledge
of the product range and characteristics manufactured and/or
assembled in this system. In this context, the main challenge in
modelling and analysis is now not only to cope with single
products, a limited product range or existing product families,
but also to be able to analyze and to compare products to define
new product families. It can be observed that classical existing
product families are regrouped in function of clients or features.
However, assembly oriented product families are hardly to find.
On the product family level, products differ mainly in two
main characteristics: (i) the number of components and (ii) the
type of components (e.g. mechanical, electrical, electronical).
Classical methodologies considering mainly single products
or solitary, already existing product families analyze the
product structure on a physical level (components level) which
causes difficulties regarding an efficient definition and
comparison of different product families. Addressing this
Procedia CIRP 102 (2021) 7–12
2212-8271 © 2021 The Authors. Published by Elsevier B.V.
This is an open access article under the CC BY-NC-ND license (https://creativecommons.org/licenses/by-nc-nd/4.0)
Peer-review under responsibility of the scientific committee of the 18th CIRP Conference on Modeling of Machining Operation.
10.1016/j.procir.2021.09.002
© 2021 The Authors. Published by Elsevier B.V.
This is an open access article under the CC BY-NC-ND license (https://creativecommons.org/licenses/by-nc-nd/4.0)
Peer-review under responsibility of the scientic committee of the 18th CIRP Conference on Modeling of Machining Operation.
Available online at www.sciencedirect.com
ScienceDirect
Procedia CIRP 00 (2019) 000–000
www.elsevier.com/locate/procedia
2212-8271 © 2021 The Authors. Published by ELSEVIER B.V.
This is an open access article under the CC BY-NC-ND license (https://creativecommons.org/licenses/by-nc-nd/4.0)
Peer-review under responsibility of the scientific committee of the 18th CIRP Conference on Modeling of Machining Operation
18th CIRP Conference on Modeling of Machining Operations
Experimental and FEM analysis of dry and cryogenic turning of hardened
steel 100Cr6 using CBN Wiper tools
G. Ortiz-de-Zaratea,*, D. Sorianoa, A. Madariagaa, A. Garaya, I. Rodrigueza, P.J. Arrazolaa
a Mondragon Unibertsitatea, Faculty of Engineering, Loramendi 4, Arrasate-Mondragón, 20500, Spain
* Corresponding author. Tel.: +34 943 79 47 00; fax: +34 943 79 15 36. E-mail address: gortizdezarate@mondragon.edu
Abstract
Employing cutting fluids in machining processes, especially for difficult-to-cut materials, improves machinability through prolonged tool life,
improves surface integrity and chip evacuation. However, like oil and water-based cutting fluids are hazardous to the environment and workers'
health, alternative solutions are required. Liquid Nitrogen (LN2) is a cryogenic fluid that can be an option due to its low boiling point (-197ºC)
and the fact it exists in the atmosphere at room conditions. Nevertheless, the feasibility of cryogenic cooling techniques in machining is not fully
understood; this is why the Finite Element Method (FEM) could give an insight into the phenomena happening on the tool-chip/workpiece
interface. This research aims to compare fundamental and industrial outputs when turning hardened steel 100Cr6 using Cubic Boron Nitride
(CBN) inserts with wiper geometry in dry conditions and with cryogenic cooling. For this purpose, turning experimental tests were performed in
both cooling conditions varying the cutting speed (150-550 m/min). Machining forces were measured during the tests, and then tool wear,
microstructural damage, and residual stresses of the workpiece were characterised. A nose turning (3D) FEM model was also developed to
understand the influence of cooling strategy on the outputs measured experimentally.
© 2021 The Authors. Published by ELSEVIER B.V.
This is an open access article under the CC BY-NC-ND license (https://creativecommons.org/licenses/by-nc-nd/4.0)
Peer-review under responsibility of the scientific committee of the 18th CIRP Conference on Modeling of Machining Operation
Keywords: Cryogenic; FEM; Nose turning; 100Cr6; CBN
1. Introduction
Hardened 100Cr6 bearing steel (62 HRC) is a difficult-to-
cut alloy due to its high hardness. Machining of hard materials
increases cutting forces and accelerates tool wear [1]. As
cutting forces increase, so does the energy required to cut the
material. Most of the energy in the cutting process is converted
into heat, which weakens the tool material's interparticle
bonding, further increasing tool wear and worsening the surface
integrity of the resultant machined part [1].
To avoid such machinability and also surface integrity
problems, extra-hard tool materials, such as Cubic Boron
Nitride (CBN) tools, have been used by various researchers
[2,3]. Another solution employed when machining difficult-to-
cut materials relies on the use of cutting fluids. Oil and water-
based cutting fluids show a significant improvement in tool life
and surface integrity due to their capability of extracting heat
and lubricating the cutting zone [4]. On the downside, these
fluids can create several environmental and health problems,
such as water pollution and soil contamination at disposal, and
generation of fumes, which can be hazardous for workers [5].
Extensive research is being conducted in recent decades to
develop new machines and methods to reduce, eliminate, or
replace conventional cutting fluids. The use of "sub-zero"
cooling by applying liquified gases to the cutting zone has been
especially attractive for researchers over other techniques such
as Minimum Quantity Lubrication (MQL) [6].
Liquid carbon dioxide (LCO2) and liquid nitrogen (LN2)
have been researched as sustainable and high-performance
cooling alternatives to conventional cutting fluids. The reason
for this is the low temperature of these gases and their
potentially superior heat extraction capacity compared to
traditional cutting fluids. Additionally, these alternative cooling
systems also lead to the elimination of hazards at the workplace
Available online at www.sciencedirect.com
ScienceDirect
Procedia CIRP 00 (2019) 000–000
www.elsevier.com/locate/procedia
2212-8271 © 2021 The Authors. Published by ELSEVIER B.V.
This is an open access article under the CC BY-NC-ND license (https://creativecommons.org/licenses/by-nc-nd/4.0)
Peer-review under responsibility of the scientific committee of the 18th CIRP Conference on Modeling of Machining Operation
18th CIRP Conference on Modeling of Machining Operations
Experimental and FEM analysis of dry and cryogenic turning of hardened
steel 100Cr6 using CBN Wiper tools
G. Ortiz-de-Zaratea,*, D. Sorianoa, A. Madariagaa, A. Garaya, I. Rodrigueza, P.J. Arrazolaa
a Mondragon Unibertsitatea, Faculty of Engineering, Loramendi 4, Arrasate-Mondragón, 20500, Spain
* Corresponding author. Tel.: +34 943 79 47 00; fax: +34 943 79 15 36. E-mail address: gortizdezarate@mondragon.edu
Abstract
Employing cutting fluids in machining processes, especially for difficult-to-cut materials, improves machinability through prolonged tool life,
improves surface integrity and chip evacuation. However, like oil and water-based cutting fluids are hazardous to the environment and workers'
health, alternative solutions are required. Liquid Nitrogen (LN2) is a cryogenic fluid that can be an option due to its low boiling point (-197ºC)
and the fact it exists in the atmosphere at room conditions. Nevertheless, the feasibility of cryogenic cooling techniques in machining is not fully
understood; this is why the Finite Element Method (FEM) could give an insight into the phenomena happening on the tool-chip/workpiece
interface. This research aims to compare fundamental and industrial outputs when turning hardened steel 100Cr6 using Cubic Boron Nitride
(CBN) inserts with wiper geometry in dry conditions and with cryogenic cooling. For this purpose, turning experimental tests were performed in
both cooling conditions varying the cutting speed (150-550 m/min). Machining forces were measured during the tests, and then tool wear,
microstructural damage, and residual stresses of the workpiece were characterised. A nose turning (3D) FEM model was also developed to
understand the influence of cooling strategy on the outputs measured experimentally.
© 2021 The Authors. Published by ELSEVIER B.V.
This is an open access article under the CC BY-NC-ND license (https://creativecommons.org/licenses/by-nc-nd/4.0)
Peer-review under responsibility of the scientific committee of the 18th CIRP Conference on Modeling of Machining Operation
Keywords: Cryogenic; FEM; Nose turning; 100Cr6; CBN
1. Introduction
Hardened 100Cr6 bearing steel (62 HRC) is a difficult-to-
cut alloy due to its high hardness. Machining of hard materials
increases cutting forces and accelerates tool wear [1]. As
cutting forces increase, so does the energy required to cut the
material. Most of the energy in the cutting process is converted
into heat, which weakens the tool material's interparticle
bonding, further increasing tool wear and worsening the surface
integrity of the resultant machined part [1].
To avoid such machinability and also surface integrity
problems, extra-hard tool materials, such as Cubic Boron
Nitride (CBN) tools, have been used by various researchers
[2,3]. Another solution employed when machining difficult-to-
cut materials relies on the use of cutting fluids. Oil and water-
based cutting fluids show a significant improvement in tool life
and surface integrity due to their capability of extracting heat
and lubricating the cutting zone [4]. On the downside, these
fluids can create several environmental and health problems,
such as water pollution and soil contamination at disposal, and
generation of fumes, which can be hazardous for workers [5].
Extensive research is being conducted in recent decades to
develop new machines and methods to reduce, eliminate, or
replace conventional cutting fluids. The use of "sub-zero"
cooling by applying liquified gases to the cutting zone has been
especially attractive for researchers over other techniques such
as Minimum Quantity Lubrication (MQL) [6].
Liquid carbon dioxide (LCO2) and liquid nitrogen (LN2)
have been researched as sustainable and high-performance
cooling alternatives to conventional cutting fluids. The reason
for this is the low temperature of these gases and their
potentially superior heat extraction capacity compared to
traditional cutting fluids. Additionally, these alternative cooling
systems also lead to the elimination of hazards at the workplace
8 G. Ortiz-de-Zarate et al. / Procedia CIRP 102 (2021) 7–12
2 G. Ortiz-de-Zarate / Procedia CIRP 00 (2019) 000–000
since the presence of these gases is non-hazardous (they exist
in the atmosphere).
Nevertheless, the feasibility of the cryogenic cooling
technique is still not fully understood since, as opposed to
conventional cooling, cryogenic jet cooling is highly sensitive
to the position of the nozzles with respect to the cutting tool and
workpiece [7,8]. In some cases, the LN2 can overcool the
workpiece or fail to impinge on the tool-chip interface
correctly, leading to adverse results in tool life and surface
integrity [9,10].
Finite Element Method (FEM) has been used in many
materials and processes to analyse and understand the
behaviour of difficult-to-cut materials in machining processes.
Nevertheless, there is currently a very small amount of
published works that consider the cryogenic coolant effect.
Among them, Pusavec et al. [11] numerically predicted the
phase and the surface heat transfer coefficient of LN2 in
machining. Once the heat transfer models were verified, they
were applied to different surfaces of the cutting tools (rake face
and flank face) and in the surface of the workpiece prior to
machining. These tests were validated via orthogonal cutting of
Inconel 718 under dry and cryogenic conditions. Salame et al.
[12] combined FEM and Computational Fluid Dynamics (CFD)
modelling for understanding the effect of different nozzle
positions on the physical parameters of the fluid coming in
contact with the tool-workpiece and the subsequent impact on
cutting temperatures and tool wear. They modelled the
impingement of the LCO2 "sub-zero" jet and the stress,
temperature distribution, and cutting forces in orthogonal
cutting of Ti-6Al-4V. However, the model was not validated
experimentally.
The literature review evidences the necessity of gaining
knowledge on the phenomena involved in the lubricant-
tool/workpiece interface to understand the cooling role in the
generation of tool wear and surface integrity.
This research aims to analyse the influence of cooling (dry
and LN2) and cutting speed in fundamental (forces) and
industrial outcomes (tool wear, microstructural damage, and
residual stresses) while turning hardened steel 100Cr6 with
CBN tools. To analyse the physical phenomena that occur in
the lubricant-tool/workpiece interface that explains the outputs
obtained experimentally, a 3D FEM model for dry and LN2
cooling conditions was developed.
2. Experimental tests
The experimental tests were carried out in a CNC Fagor
8070 T horizontal lathe, using a PCLN L 2020 K12 tool holder
and CNGA120408S01030AWH7015 CBN wiper coated insert
(PVD TiN coating thickness of 4±0.5 μm). The workpiece was
a bar (Ø89 x 450 mm) of 100Cr6 hardened steel with 62 HRC.
Regarding the cutting conditions, the feed and depth of cut
were fixed, while the cutting speed and lubrication were varied.
In dry condition, the cutting speed influence was analysed
every 50 m/min from 300 to 550 m/min, while with cryogenic
cooling, the analysis was extended from 150 until 550 m/min
(see Table 1). Two repetitions were done for each cutting
condition, machining 30 mm of the bar in the longitudinal
direction per test.
Table 1. Experimental cutting conditions.
Cutting condition parameter
Dry
LN
2
Cutting speed, v
c
(m/min)
300-550
150-550
Feed, f (mm/rev)
0.2
Depth of cut, a
p
(mm)
0.2
The LN2 system supplied cryogenic coolant to the shim of
the tool holder (see Fig. 1). This shim had a pilot hole to direct
the LN2 flow to the clearance face of the insert. The aim was to
cool the newly machined surface with the LN2 , while keeping
the tool cool. The cooling system had a phase separator which
ensured that the nitrogen was supplied in liquid phase to the
cutting zone. The setup also incorporated a Kistler 9121
dynamometer to measure the forces in the three directions. The
same setup was used for dry tests, removing the LN2 system.
Fig. 1. Experimental setup for LN2.
After each test, the tool flank wear was measured with an
optical amplifier. The microstructural damage of the machined
surface was also observed using an optical microscope. For
that, cutups in the cutting and feed directions were done, and
samples were prepared following conventional metallographic
techniques. Additionally, the surface residual stresses were
measured by X-ray diffraction in the cutting and feed directions
for a single cutting condition (vc = 300 m/min, f = 0.2 mm/rev,
and ap = 0.2 mm) under dry and LN2 cooling conditions. A
Bruker D8 advance diffractometer was employed for this
purpose. The radiation used was CrKα1 with a voltage of 40 kV
and a current of 4 mA. The (2 1 1) diffraction plane was chosen
for the measurements. A round collimator (2 mm diameter) on
the incident beam was used. Measurements were carried out in
Ω mode with 14 psi inclinations, ranging from -50º to 50º.
Diffraction peaks were fitted with a Pearson VII function that
is necessary for eliminating errors from varying blending and
defocusing of the Kα doublet diffraction peak [13]. The
diffraction elastic constants used in the measurements were:
S1 = -1.271 E-6 (MPa-1) and 1
2S2 = 5.811 E-6 (MPa-1).
3. Finite Element modelling
Deform 3D FEM commercial software was used to develop
a 3D nose turning model for dry and cryogenic cooling
conditions (see Fig. 2). The software incorporates a lagrangian
implicit formulation with remeshing to avoid mesh distortion.
LN
2
tank
N
2
gas
LN
2
liquid
100Cr6
workpiece
LN
2
nozzle
Phase separator
Kistler 9121
CBN insert
Pilot hole
G. Ortiz-de-Zarate / Procedia CIRP 00 (2019) 000–000 3
Fig. 2. Boundary conditions of the FEM nose turning model.
The model consists of a rigid tool and an elastoplastic
workpiece meshed with 50,000 and 80,000 elements,
respectively, with a minimum element size in the workpiece of
10 μm. The remeshing criteria was established to locate the
minimum element size in the areas with the higher strain and
strain rate, i.e., in the shear regions, including the tool-chip
interface, and machined surface.
The tool geometry was characterised using the Alicona
IFG4 optical profilometer. First, the 3D geometry of the wiper
insert was obtained (Fig. 3a). Then, the rake and relief angles,
cutting edge radius, and chamfer geometry were measured
(Fig. 3b and Table 2). Using SolidWorks 2020, the tool was
modelled (Fig. 3c), and it was imported into the FEM software
in STL format. Finally, the tool orientation with respect to the
workpiece was corrected in the FEM model to match the tool
holder’s angles (-6 º axial and radial directions).
Table 2. Wiper insert geometry.
Parameter name
Value
Cutting edge radius, rβ (μm)
33
True negative bevel length, bγ (μm)
137
Angle of the negative bevel γb (º)
30
Clearance angle, α (º)
0
Rake angle, γ (º)
0
Regarding the material properties, the thermo-viscoplastic
behaviour of the workpiece was modelled with Johnson Cook's
law. The parameters were obtained from Ramesh and Melkote
[14]. This material undergoes a softening instead of hardening
at high strains (strain softening) [15]. Therefore, a limitation to
the yield stress was applied when the plastic strain was greater
than one by programing a subroutine in Fortran language. The
tool material properties of the CBN were obtained from the
library of the software.
Fig. 3. (a) 3D profile of the tool, (b) tool geometry parameters and
(c) SolidWorks model of the tool.
To model the tool-chip contact, the sticking-sliding model
was used since it was demonstrated to be the most
representative of machining [16]. It was selected the same
friction coefficient µ value for both lubrication conditions,
since as presented by Hong et al. [17], the application of LN2
on the clearance face does not produce significant variation of
friction with respect to dry machining. They observed that µ
varied between 0.38 (LN2) and 0.48 (dry) at vc = 90 m/min.
This could be related to the high pressures in the contact region
that does not let the lubrication to reach the cutting zone. At
higher cutting speeds, lower friction would be expected due to
the increase of the sliding velocity. Therefore, it was selected
μ = 0.3 and m = 1 for the cutting conditions used in this study.
The difference between dry and LN2 cooling simulations lies
in the boundary conditions applied to the model (see Fig. 2). In
both models, the workpiece's movement was restricted in the
three directions, and a thermal exchange with the workpiece
and with the environment was established. The air convection
coefficient used was 20 W/m2K with a room temperature of
20ºC [18]. In the LN2 cooling simulation, a heat exchange
window was included in the flank face of the tool (to reproduce
the experimental setup) that moves at the cutting speed (see
Fig. 2). In this window, the temperature was -197ºC and the
convection coefficient was 5,000 kW/m2K [18].
The simulated cutting conditions were vc = 300 m/min,
f = 0.2 mm, and ap = 0.2 mm since it is the highest cutting
speed in which the tool flank wear is low for both lubrication
conditions (see Fig. 4). These conditions were chosen because
the purpose of this study is not to analyse the influence of tool
wear but the phenomena involved in the lubrication.
4. Results and discussion
4.1. Fundamental variables and tool wear
Fig. 4a shows the influence of cutting speed on maximum
experimental cutting (Fc), feed (Ff), and passive (Fp) forces.
Fig. 4b presents the tool flank wear for the different cutting
conditions tested. In cryogenic cooling, two clear different tool
wear regions were observed that match with the force trends.
On the one hand, below 300 m/min, homogenous tool wear was
reported, and consequently, the cutting speed produces no
significant variation in forces, as also observed by [19]. On the
other hand, above 300 m/min, a drastic increase of tool wear
was observed for the LN2 cooling conditions that produced the
sudden increase of the forces. This could be related to the
potential embrittlement of the CBN tool produced by the
thermal shock of the cryogenic cooling at high cutting speeds
that promotes tool wear (see Fig. 4b,c). This was also observed
in the literature on carbide tools [20].
In dry turning, the increase of the cutting speed reduces
forces due to the rise of the thermal softening and the decrease
of chip thickness, as expected from the literature [21,22]. The
tool wear appears not to be significantly affected by the cutting
speed (see Fig. 4b,c).
As abovementioned, the comparison between experimental
and FEM model results was made for vc = 300 m/min to avoid
the drastic tool wear region. The forces of the FEM model were
obtained in the thermal steady-state condition, with a machined
a) b)
(a)
a) b)
(c)
γ
α
γ
b
(b)
rβ
Heat exchange window
for LN
2
simulation
Heat exchange with
environment
Heat exchange with
workpiece
V
x
= V
y
= V
z
= 0 XY
Z
v
f
G. Ortiz-de-Zarate et al. / Procedia CIRP 102 (2021) 7–12 9
2 G. Ortiz-de-Zarate / Procedia CIRP 00 (2019) 000–000
since the presence of these gases is non-hazardous (they exist
in the atmosphere).
Nevertheless, the feasibility of the cryogenic cooling
technique is still not fully understood since, as opposed to
conventional cooling, cryogenic jet cooling is highly sensitive
to the position of the nozzles with respect to the cutting tool and
workpiece [7,8]. In some cases, the LN2 can overcool the
workpiece or fail to impinge on the tool-chip interface
correctly, leading to adverse results in tool life and surface
integrity [9,10].
Finite Element Method (FEM) has been used in many
materials and processes to analyse and understand the
behaviour of difficult-to-cut materials in machining processes.
Nevertheless, there is currently a very small amount of
published works that consider the cryogenic coolant effect.
Among them, Pusavec et al. [11] numerically predicted the
phase and the surface heat transfer coefficient of LN2 in
machining. Once the heat transfer models were verified, they
were applied to different surfaces of the cutting tools (rake face
and flank face) and in the surface of the workpiece prior to
machining. These tests were validated via orthogonal cutting of
Inconel 718 under dry and cryogenic conditions. Salame et al.
[12] combined FEM and Computational Fluid Dynamics (CFD)
modelling for understanding the effect of different nozzle
positions on the physical parameters of the fluid coming in
contact with the tool-workpiece and the subsequent impact on
cutting temperatures and tool wear. They modelled the
impingement of the LCO2 "sub-zero" jet and the stress,
temperature distribution, and cutting forces in orthogonal
cutting of Ti-6Al-4V. However, the model was not validated
experimentally.
The literature review evidences the necessity of gaining
knowledge on the phenomena involved in the lubricant-
tool/workpiece interface to understand the cooling role in the
generation of tool wear and surface integrity.
This research aims to analyse the influence of cooling (dry
and LN2) and cutting speed in fundamental (forces) and
industrial outcomes (tool wear, microstructural damage, and
residual stresses) while turning hardened steel 100Cr6 with
CBN tools. To analyse the physical phenomena that occur in
the lubricant-tool/workpiece interface that explains the outputs
obtained experimentally, a 3D FEM model for dry and LN2
cooling conditions was developed.
2. Experimental tests
The experimental tests were carried out in a CNC Fagor
8070 T horizontal lathe, using a PCLN L 2020 K12 tool holder
and CNGA120408S01030AWH7015 CBN wiper coated insert
(PVD TiN coating thickness of 4±0.5 μm). The workpiece was
a bar (Ø89 x 450 mm) of 100Cr6 hardened steel with 62 HRC.
Regarding the cutting conditions, the feed and depth of cut
were fixed, while the cutting speed and lubrication were varied.
In dry condition, the cutting speed influence was analysed
every 50 m/min from 300 to 550 m/min, while with cryogenic
cooling, the analysis was extended from 150 until 550 m/min
(see Table 1). Two repetitions were done for each cutting
condition, machining 30 mm of the bar in the longitudinal
direction per test.
Table 1. Experimental cutting conditions.
Cutting condition parameter
Dry
LN2
Cutting speed, vc (m/min)
300-550
150-550
Feed, f (mm/rev)
0.2
Depth of cut, ap (mm)
0.2
The LN2 system supplied cryogenic coolant to the shim of
the tool holder (see Fig. 1). This shim had a pilot hole to direct
the LN2 flow to the clearance face of the insert. The aim was to
cool the newly machined surface with the LN2 , while keeping
the tool cool. The cooling system had a phase separator which
ensured that the nitrogen was supplied in liquid phase to the
cutting zone. The setup also incorporated a Kistler 9121
dynamometer to measure the forces in the three directions. The
same setup was used for dry tests, removing the LN2 system.
Fig. 1. Experimental setup for LN2.
After each test, the tool flank wear was measured with an
optical amplifier. The microstructural damage of the machined
surface was also observed using an optical microscope. For
that, cutups in the cutting and feed directions were done, and
samples were prepared following conventional metallographic
techniques. Additionally, the surface residual stresses were
measured by X-ray diffraction in the cutting and feed directions
for a single cutting condition (vc = 300 m/min, f = 0.2 mm/rev,
and ap = 0.2 mm) under dry and LN2 cooling conditions. A
Bruker D8 advance diffractometer was employed for this
purpose. The radiation used was CrKα1 with a voltage of 40 kV
and a current of 4 mA. The (2 1 1) diffraction plane was chosen
for the measurements. A round collimator (2 mm diameter) on
the incident beam was used. Measurements were carried out in
Ω mode with 14 psi inclinations, ranging from -50º to 50º.
Diffraction peaks were fitted with a Pearson VII function that
is necessary for eliminating errors from varying blending and
defocusing of the Kα doublet diffraction peak [13]. The
diffraction elastic constants used in the measurements were:
S1 = -1.271 E-6 (MPa-1) and 1
2S2 = 5.811 E-6 (MPa-1).
3. Finite Element modelling
Deform 3D FEM commercial software was used to develop
a 3D nose turning model for dry and cryogenic cooling
conditions (see Fig. 2). The software incorporates a lagrangian
implicit formulation with remeshing to avoid mesh distortion.
LN
2
tank
N
2
gas
LN
2
liquid
100Cr6
workpiece
LN
2
nozzle
Phase separator
Kistler 9121
CBN insert
Pilot hole
G. Ortiz-de-Zarate / Procedia CIRP 00 (2019) 000–000 3
Fig. 2. Boundary conditions of the FEM nose turning model.
The model consists of a rigid tool and an elastoplastic
workpiece meshed with 50,000 and 80,000 elements,
respectively, with a minimum element size in the workpiece of
10 μm. The remeshing criteria was established to locate the
minimum element size in the areas with the higher strain and
strain rate, i.e., in the shear regions, including the tool-chip
interface, and machined surface.
The tool geometry was characterised using the Alicona
IFG4 optical profilometer. First, the 3D geometry of the wiper
insert was obtained (Fig. 3a). Then, the rake and relief angles,
cutting edge radius, and chamfer geometry were measured
(Fig. 3b and Table 2). Using SolidWorks 2020, the tool was
modelled (Fig. 3c), and it was imported into the FEM software
in STL format. Finally, the tool orientation with respect to the
workpiece was corrected in the FEM model to match the tool
holder’s angles (-6 º axial and radial directions).
Table 2. Wiper insert geometry.
Parameter name
Value
Cutting edge radius, r
β
(μm)
33
True negative bevel length, b
γ
(μm)
137
Angle of the negative bevel γ
b
(º)
30
Clearance angle, α (º)
0
Rake angle, γ (º)
0
Regarding the material properties, the thermo-viscoplastic
behaviour of the workpiece was modelled with Johnson Cook's
law. The parameters were obtained from Ramesh and Melkote
[14]. This material undergoes a softening instead of hardening
at high strains (strain softening) [15]. Therefore, a limitation to
the yield stress was applied when the plastic strain was greater
than one by programing a subroutine in Fortran language. The
tool material properties of the CBN were obtained from the
library of the software.
Fig. 3. (a) 3D profile of the tool, (b) tool geometry parameters and
(c) SolidWorks model of the tool.
To model the tool-chip contact, the sticking-sliding model
was used since it was demonstrated to be the most
representative of machining [16]. It was selected the same
friction coefficient µ value for both lubrication conditions,
since as presented by Hong et al. [17], the application of LN2
on the clearance face does not produce significant variation of
friction with respect to dry machining. They observed that µ
varied between 0.38 (LN2) and 0.48 (dry) at vc = 90 m/min.
This could be related to the high pressures in the contact region
that does not let the lubrication to reach the cutting zone. At
higher cutting speeds, lower friction would be expected due to
the increase of the sliding velocity. Therefore, it was selected
μ = 0.3 and m = 1 for the cutting conditions used in this study.
The difference between dry and LN2 cooling simulations lies
in the boundary conditions applied to the model (see Fig. 2). In
both models, the workpiece's movement was restricted in the
three directions, and a thermal exchange with the workpiece
and with the environment was established. The air convection
coefficient used was 20 W/m2K with a room temperature of
20ºC [18]. In the LN2 cooling simulation, a heat exchange
window was included in the flank face of the tool (to reproduce
the experimental setup) that moves at the cutting speed (see
Fig. 2). In this window, the temperature was -197ºC and the
convection coefficient was 5,000 kW/m2K [18].
The simulated cutting conditions were vc = 300 m/min,
f = 0.2 mm, and ap = 0.2 mm since it is the highest cutting
speed in which the tool flank wear is low for both lubrication
conditions (see Fig. 4). These conditions were chosen because
the purpose of this study is not to analyse the influence of tool
wear but the phenomena involved in the lubrication.
4. Results and discussion
4.1. Fundamental variables and tool wear
Fig. 4a shows the influence of cutting speed on maximum
experimental cutting (Fc), feed (Ff), and passive (Fp) forces.
Fig. 4b presents the tool flank wear for the different cutting
conditions tested. In cryogenic cooling, two clear different tool
wear regions were observed that match with the force trends.
On the one hand, below 300 m/min, homogenous tool wear was
reported, and consequently, the cutting speed produces no
significant variation in forces, as also observed by [19]. On the
other hand, above 300 m/min, a drastic increase of tool wear
was observed for the LN2 cooling conditions that produced the
sudden increase of the forces. This could be related to the
potential embrittlement of the CBN tool produced by the
thermal shock of the cryogenic cooling at high cutting speeds
that promotes tool wear (see Fig. 4b,c). This was also observed
in the literature on carbide tools [20].
In dry turning, the increase of the cutting speed reduces
forces due to the rise of the thermal softening and the decrease
of chip thickness, as expected from the literature [21,22]. The
tool wear appears not to be significantly affected by the cutting
speed (see Fig. 4b,c).
As abovementioned, the comparison between experimental
and FEM model results was made for vc = 300 m/min to avoid
the drastic tool wear region. The forces of the FEM model were
obtained in the thermal steady-state condition, with a machined
a) b)
(a)
a) b)
(c)
γ
α
γ
b
(b)
rβ
Heat exchange window
for LN
2
simulation
Heat exchange with
environment
Heat exchange with
workpiece
V
x
= V
y
= V
z
= 0 XY
Z
v
f
10 G. Ortiz-de-Zarate et al. / Procedia CIRP 102 (2021) 7–12
4 G. Ortiz-de-Zarate / Procedia CIRP 00 (2019) 000–000
Fig. 4. Influence of cutting speed in (a) maximum forces (uncertainty of 10%)
and (b) tool flank wear. (c) Tool flank wear for vc = 300 m/min and
550 m/min in dry and LN2.
distance of 2 mm. The FEM model shows the same trends as
the experimental results since both reported slightly higher
forces for the cryogenic cooling conditions than for dry
machining (see Fig. 5). As for the quantitative values, the
model predicts lower values than the experimental ones. This
might occur because of three main reasons:
• The analysed forces are the maximum, which means that
they were obtained at the end of the test, where the tool wear
was higher. Although the tool wear for 300 m/min was low
(0.13-0.17 mm), it might have influenced the forces.
• It is difficult to ensure that the heat exchange window
represents precisely the behaviour of the LN2 when it
impacts the workpiece. In the experimental tests, there may
be splashes that could vary the cooling of the tool/workpiece
that are not modelled in the FEM approach. The numerical
model assumes a very localised and homogeneous
impingement of the LN2 jet (see the cooled region in
temperature results of Fig. 6).
• The flow stress and friction parameters were not
experimentally characterised. This means that the material
parameters might not represent correctly the
thermomechanical behaviour of the workpiece material used
in the experimental tests.
The FEM model was also consistent in the influence of the
cooling in the temperatures. The most significant difference
was observed in the temperature gradient along the machined
surface (see Fig. 6a,b). Dry condition shows a slow and
progressive cooling of the machined surface, while cryogenic
results in fast cooling with a high temperature gradient between
the tertiary shear zone and the machined surface.
Fig. 5. Experimental and FEM model results of forces (vc = 300 m/min,
f = 0.2 mm, ap = 0.2 mm).
In the primary and secondary shear zones, the temperature
differences were significantly low. As expected, dry condition
shows higher temperatures, which is consistent with the force
results. However, the differences are not as significant as might
be expected. This could be due to the fact that the LN2 only
cools the tool on the flank face, so it reaches the contact area
by conduction, reducing its efficiency in cooling the contact
area (see Fig. 6c,d). That produces that the tool's maximum
temperature is similar in both conditions. The main difference
is the higher thermal gradient in cryogenic conditions. This
might produce that at higher cutting speeds, where higher
contact temperatures are expected, the thermal shock becomes
so significant that it could promote drastic tool wear as
observed experimentally.
Fig. 6. Workpiece (a,b) and tool (c,d) temperatures in dry (a,c) and
cryogenic (b,d) conditions (vc = 300 m/min, f = 0.2 mm, ap = 0.2 mm).
4.2. Microstructural damage
The microstructural damage was analysed in the
experimental tests observing samples prepared from cutups in
the cutting and feed directions (see Fig. 7). The main defects
were the presence of a thermally affected layer, white layer,
and adhered material. All the machined surfaces had a
thermally affected layer in the cutting direction, independent of
the cutting speed or lubrication conditions (see Fig. 8).
The cutups in feed direction show step-tool with adhered
material only in the surfaces machined with LN2 at cutting
speeds above 400 m/min (see Fig. 7e-h). The frequency of the
step-tool matches the feed rate. This could be explained by the
drastic tool wear observed at the highest cutting speeds.
Moreover, the increase in cutting speed seems to reduce the
height of the adhered material, from 16.9 μm for
vc = 400 m/min to 2.6 μm for vc = 550 m/min.
(a) (b)
Temperature (ºC)
800
688
575
463
350
238
125
12
-100
(c) (d)
0
100
200
300
400
500
600
700
800
Maximum forces (N)
Dry
LN
2
F
c
0
100
200
300
400
500
600
700
800
150 200 250 300 350 400 450 500 550
Maximum forces (N)
Cutting speed (m/min)
Seco_Fx Seco_Fy Seco_Fz
LN2_Fx LN2_Fy LN2_Fz
0
100
200
300
400
500
600
700
800
150 200 250 300 350 400 450 500 550
Maximum forces (N)
Cutting speed (m/min)
Seco_Fx Seco_Fy Seco_Fz
LN2_Fx LN2_Fy LN2_Fz
0
100
200
300
400
500
600
700
800
150 200 250 300 350 400 450 500 550
Maximum forces (N)
Cutting speed (m/min)
Seco_Fx Seco_Fy Seco_Fz
LN2_Fx LN2_Fy LN2_Fz
F
p
F
f
0
0.2
0.4
0.6
0.8
1
1.2
1.4
1.6
1.8
150 200 250 300 350 400 450 500 550
Flank wear, V
b
(mm)
Cutting speed (m/min)
Insert breakage
Drastic tool wear
Homogeneous
tool wear
(a)
(b)
0
0.2
0.4
0.6
0.8
1
1.2
1.4
1.6
1.8
150 200 250 300 350 400 450 500 550
Flank wear, V
b
(mm)
Cutting speed (m/min)
Seco LN2
0
0.2
0.4
0.6
0.8
1
1.2
1.4
1.6
1.8
150 200 250 300 350 400 450 500 550
Flank wear, V
b
(mm)
Cutting speed (m/min)
Seco LN2
Dry
LN
2
1 mm
(c)
1 mm1 mm
1 mm
Dry
v
c
= 300 m/min
LN
2
v
c
= 300 m/min
Dry
v
c
= 550 m/min
LN
2
v
c
= 550 m/min
0
100
200
300
Forces (N)
0
100
200
300
Forces (N)
Series1
Series2
Series3
Series4
0
100
200
300
Forces (N)
Series1
Series2
Series3
Series4
Dry LN
2
FcFp
Ff
FEM model
Experimental
G. Ortiz-de-Zarate / Procedia CIRP 00 (2019) 000–000 5
Fig. 7. Cutups in (a-d) cutting and (e-h) feed direction.
Moreover, when machining under cryogenic cooling white
layer was observed in most of the cutting speeds, while in dry,
it was only found at 400 m/min (see Fig. 8). Generally, the
white layer appears in severe cutting conditions (high cutting
speed and increased flank wear) when there are high
temperatures in the tertiary shear zone. Nevertheless, in LN2
cooling conditions it can be observed for all cutting speeds
(except for 350 and 550 m/min) and even for low tool wear
conditions. Kaynak et al. [23] and Pu et al. [24] also observed
more significant microstructural damage for LN2 than for dry
for different materials. Furthermore, a similar white layer was
obtained for all the cutting conditions tested in LN2 cooling in
this work. Hence, tool wear and cutting speed seem not to
significantly affect the white layer depth (see Fig. 8).
To find an explanation for this phenomenon, the strains and
temperatures in the machined surface using the FEM model
were analysed. The tertiary shear zone temperatures appear to
be similar for both cooling conditions (see Fig. 9a,b). Although
with LN2 the cooling effect is drastically higher, it seems that
the fluid did not reach the tertiary shear zone due to the high
contact pressures of the region. Thus, a similar thermally
affected layer would be expected in the machined surface as
observed experimentally. Nevertheless, the white layer could
not be explained from the surface temperature point of view,
and neither from the surface strain, since similar values are
reported in the FEM model (see Fig. 9). The only difference
Fig. 8. Influence of cutting speed and lubrication in microstructural damage.
Fig. 9. Temperature (a,b) and plastic strain (c,d) obtained in the machined
surface in dry (a,c) and cryogenic (b,d) conditions (vc = 300 m/min,
f = 0.2 mm, ap = 0.2 mm).
between lubrication conditions is the higher thermal gradient
along the machined surface in LN2 condition that may promote
the white layer generation, but more research is needed to
confirm this claim.
4.3. Residual stresses
Fig. 10 shows the average values of the residual stresses
measured at the surface of the specimens turned using different
cooling strategies. Surface residual stresses were tensile in both
the cutting (431 ± 130 MPa for dry and 529 ± 84 MPa for LN2)
and feed direction (123 ± 38 MPa for dry and 90 ± 84 MPa for
LN2). It is well known in the field that machining-induced
residual stresses have two primary sources: i) thermal loads
induce tensile stresses near the surface, and ii) machining
forces produce more compressive residual stresses. Therefore,
the thermal loads generated during the cutting processes had a
greater impact on the surface residual stresses than the
machining forces for the tested conditions and material. It
should be noted that tensile residual stresses were more tensile
in the cutting direction than in the feed direction, which is also
in agreement with the literature results.
No significant variations were observed between the cooling
strategies since differences are within deviation bars. This
finding is in agreement with force measurements and
predictions, where no significant differences were detected.
Furthermore, simulations also predicted similar contact
temperatures in the tertiary shear zone. The slightly higher
tensile residual stresses generated by LN2 in the cutting
direction could be attributed to the more rapid cooling as
observed in the FEM results. However, this observation needs
further analysis, since it is not clear that the fluid reached the
tertiary shear zone, as explained in Section 4.2.
Fig. 10. Surface residual stresses in dry and cryogenic condition
(vc = 300 m/min, f = 0.2 mm/rev, ap = 0.2 mm).
0
200
400
600
Residual stress (MPa)
0
200
400
600
Residual
stress …
Dry
LN2
Dry
LN
2
σ
1
= σ
cutting
σ
2
= σ
feed
σ
1
= σ
cutting
σ
2
= σ
feed
50 μm
(a) 14.5 μm(b)
14.8 μm14.3 μm1.6 μm14.2 μm16.2 μm
50 μm
14.0 μm1.8 μm14.6 μm
(d)
(c) 14.7 μm14.4 μm14.2 μm
50 μm
16.9 μm
201.8 μm
199.2 μm
6.5 μm
50 μm50 μm
(e) (f)
50 μm
50 μm
190.2 μm 191.1 μm
8.7 μm12.6 μm
210.5 μm209.4 μm
(g) (h)
Dry
vc= 300 m/min
LN2
vc= 300 m/min
Dry
vc= 500 m/min
LN2
vc= 500 m/min
LN2
vc= 400 m/min
LN2
vc= 450 m/min
LN2
vc= 500 m/min
LN2
vc= 550 m/min
50 μm
193.6 μm 199.3 μm
White layer
Thermally afected layer
Step-tool with
adhered material
0
5
10
15
20
25
150 200 250 300 350 400 450 500 550
Layer thickness (μm)
Cutting speed (m/min)
Series4
Series1
Series2
Series3
0
5
10
15
20
25
150 200 250 300 350 400 450 500 550
Layer thickness (μm)
Cutting speed (m/min)
Series4
Series1
Series2
Series3
Dry
LN
2
White
layer
Thermally
affected
layer
0
5
10
15
20
25
150 200 250 300 350 400 450 500 550
Layer thickness (μm)
Cutting speed (m/min)
0
5
10
15
20
25
150 200 250 300 350 400 450 500 550
Layer thickness (μm)
Cutting speed (m/min)
Series4
Series1
Series2
Series3
0
5
10
15
20
25
150 200 250 300 350 400 450 500 550
Layer thickness (μm)
Cutting speed (m/min)
Series4
Series1
Series2
Series3
Temperature (ºC)
500
-100
350
200
50
(a) (b)
(c) (d)
Plastic strain (-)
1
0
0.75
0.5
0.25
300ºC 400ºC 500ºC
800ºC
-150ºC -140ºC
800ºC
G. Ortiz-de-Zarate et al. / Procedia CIRP 102 (2021) 7–12 11
G. Ortiz-de-Zarate / Procedia CIRP 00 (2019) 000–000 5
Fig. 7. Cutups in (a-d) cutting and (e-h) feed direction.
Moreover, when machining under cryogenic cooling white
layer was observed in most of the cutting speeds, while in dry,
it was only found at 400 m/min (see Fig. 8). Generally, the
white layer appears in severe cutting conditions (high cutting
speed and increased flank wear) when there are high
temperatures in the tertiary shear zone. Nevertheless, in LN2
cooling conditions it can be observed for all cutting speeds
(except for 350 and 550 m/min) and even for low tool wear
conditions. Kaynak et al. [23] and Pu et al. [24] also observed
more significant microstructural damage for LN2 than for dry
for different materials. Furthermore, a similar white layer was
obtained for all the cutting conditions tested in LN2 cooling in
this work. Hence, tool wear and cutting speed seem not to
significantly affect the white layer depth (see Fig. 8).
To find an explanation for this phenomenon, the strains and
temperatures in the machined surface using the FEM model
were analysed. The tertiary shear zone temperatures appear to
be similar for both cooling conditions (see Fig. 9a,b). Although
with LN2 the cooling effect is drastically higher, it seems that
the fluid did not reach the tertiary shear zone due to the high
contact pressures of the region. Thus, a similar thermally
affected layer would be expected in the machined surface as
observed experimentally. Nevertheless, the white layer could
not be explained from the surface temperature point of view,
and neither from the surface strain, since similar values are
reported in the FEM model (see Fig. 9). The only difference
Fig. 8. Influence of cutting speed and lubrication in microstructural damage.
Fig. 9. Temperature (a,b) and plastic strain (c,d) obtained in the machined
surface in dry (a,c) and cryogenic (b,d) conditions (vc = 300 m/min,
f = 0.2 mm, ap = 0.2 mm).
between lubrication conditions is the higher thermal gradient
along the machined surface in LN2 condition that may promote
the white layer generation, but more research is needed to
confirm this claim.
4.3. Residual stresses
Fig. 10 shows the average values of the residual stresses
measured at the surface of the specimens turned using different
cooling strategies. Surface residual stresses were tensile in both
the cutting (431 ± 130 MPa for dry and 529 ± 84 MPa for LN2)
and feed direction (123 ± 38 MPa for dry and 90 ± 84 MPa for
LN2). It is well known in the field that machining-induced
residual stresses have two primary sources: i) thermal loads
induce tensile stresses near the surface, and ii) machining
forces produce more compressive residual stresses. Therefore,
the thermal loads generated during the cutting processes had a
greater impact on the surface residual stresses than the
machining forces for the tested conditions and material. It
should be noted that tensile residual stresses were more tensile
in the cutting direction than in the feed direction, which is also
in agreement with the literature results.
No significant variations were observed between the cooling
strategies since differences are within deviation bars. This
finding is in agreement with force measurements and
predictions, where no significant differences were detected.
Furthermore, simulations also predicted similar contact
temperatures in the tertiary shear zone. The slightly higher
tensile residual stresses generated by LN2 in the cutting
direction could be attributed to the more rapid cooling as
observed in the FEM results. However, this observation needs
further analysis, since it is not clear that the fluid reached the
tertiary shear zone, as explained in Section 4.2.
Fig. 10. Surface residual stresses in dry and cryogenic condition
(vc = 300 m/min, f = 0.2 mm/rev, ap = 0.2 mm).
0
200
400
600
Residual stress (MPa)
0
200
400
600
Residual
stress …
Dry
LN2
Dry
LN
2
σ
1
= σ
cutting
σ
2
= σ
feed
σ
1
= σ
cutting
σ
2
= σ
feed
50 μm
(a) 14.5 μm(b)
14.8 μm14.3 μm1.6 μm14.2 μm16.2 μm
50 μm
14.0 μm1.8 μm14.6 μm
(d)
(c) 14.7 μm14.4 μm14.2 μm
50 μm
16.9 μm
201.8 μm
199.2 μm
6.5 μm
50 μm50 μm
(e) (f)
50 μm
50 μm
190.2 μm 191.1 μm
8.7 μm12.6 μm
210.5 μm209.4 μm
(g) (h)
Dry
vc= 300 m/min
LN2
vc= 300 m/min
Dry
vc= 500 m/min
LN2
vc= 500 m/min
LN2
vc= 400 m/min
LN2
vc= 450 m/min
LN2
vc= 500 m/min
LN2
vc= 550 m/min
50 μm
193.6 μm 199.3 μm
White layer
Thermally afected layer
Step-tool with
adhered material
0
5
10
15
20
25
150 200 250 300 350 400 450 500 550
Layer thickness (μm)
Cutting speed (m/min)
Series4
Series1
Series2
Series3
0
5
10
15
20
25
150 200 250 300 350 400 450 500 550
Layer thickness (μm)
Cutting speed (m/min)
Series4
Series1
Series2
Series3
Dry
LN
2
White
layer
Thermally
affected
layer
0
5
10
15
20
25
150 200 250 300 350 400 450 500 550
Layer thickness (μm)
Cutting speed (m/min)
0
5
10
15
20
25
150 200 250 300 350 400 450 500 550
Layer thickness (μm)
Cutting speed (m/min)
Series4
Series1
Series2
Series3
0
5
10
15
20
25
150 200 250 300 350 400 450 500 550
Layer thickness (μm)
Cutting speed (m/min)
Series4
Series1
Series2
Series3
Temperature (ºC)
500
-100
350
200
50
(a) (b)
(c) (d)
Plastic strain (-)
1
0
0.75
0.5
0.25
300ºC 400ºC 500ºC
800ºC
-150ºC -140ºC
800ºC
12 G. Ortiz-de-Zarate et al. / Procedia CIRP 102 (2021) 7–12
6 G. Ortiz-de-Zarate / Procedia CIRP 00 (2019) 000–000
5. Summary and conclusions
This research work analysed the influence of the cutting
speed (150-550 m/min) on fundamental and surface integrity
outputs when turning hardened 100Cr6 in dry and cryogenic
cooling. A nose turning FEM model to understand the cooling
role in the chip formation process and in the generation of
surface integrity and tool wear was also developed. The main
conclusions of the research are:
• Cryogenic cooling has two different wear regions depending
on the cutting speed. Below 300 m/min cutting speed,
homogeneous wear is reported, and slightly higher forces
than dry condition are obtained. Above 300 m/min, the
drastic tool wear increase produces the sudden increase of
forces for LN2 cooling condition. The FEM model shows
similar trends when analysing the lubrication's influence at
a cutting speed of 300 m/min.
• The microstructural analysis of the machined surface shows
a similar thermally affected layer in the cutting direction for
both cooling conditions, which might be related to similar
temperatures in the tertiary shear zone, as observed in the
FEM model. White layer was reported in the cutting
direction in most cutting speeds when applying LN2 cooling,
while it was barely observed in dry cutting. Step-tool in feed
direction when turning with LN2 at high cutting speed (in
drastic tool wear region) were also observed.
• Higher tensile residual stresses were observed in the surface
for cryogenic cooling, which might be produced by the
faster cooling of this condition as observed in the FEM
results.
Acknowledgements
The authors thank NANOCRIO (KK-2016/00012) and
CRYOMACH (INNO-20182049) projects and the grant for
Education and Training of Research Staff
(PRE_2017_1_0394).
References
[1] Shihab, S. K., Khan, Z. A., Mohammad, A., Siddiquee, A. N. A review of
turning of hard steels used in bearing and automotive applications.
Production & Manufacturing Research. 2014. 2/1:24-49.
[2] Benga, G. C., Abrao, A. M. Turning of hardened 100Cr6 bearing steel with
ceramic and PCBN cutting tools. Journal of materials processing
technology. 2003. 143:237-241.
[3] de Siqueira Galoppi, G., Stipkovic Filho, M., Batalha, G. F. Hard turning
of tempered DIN 100Cr6 steel with coated and no coated CBN inserts.
Journal of materials processing technology. 2006. 179/1-3: 146-153.
[4] Brinksmeier, E., Meyer, D., Huesmann-Cordes, A. G., Herrmann, C.
Metalworking fluids—mechanisms and performance. CIRP Annals. 2015.
64/2:605-628.
[5] Pusavec, F., Krajnik, P., Kopac, J. Transitioning to sustainable production–
Part I: application on machining technologies. Journal of Cleaner
production. 2010. 18/2:174-184.
[6] Kaynak, Y., Lu, T., Jawahir, I. S. Cryogenic machining-induced surface
integrity: a review and comparison with dry, MQL, and flood-cooled
machining. Machining Science and Technology. 2014. 18/2:149-198.
[7] Lequien, P. Etude fondamentale de l’assistance cryogénique pour
application au fraisage du Ti6Al4V. 2017. Doctoral dissertation.
[8] Heep, T., Bickert, C., Abele, E. Application of carbon dioxide snow in
machining of CGI using an additively manufactured turning tool. Journal
of Manufacturing and Materials Processing. 2019. 3/1:15.
[9] Iturbe, A., Hormaetxe, E., Garay, A., Arrazola, P. J. Surface integrity
analysis when machining inconel 718 with conventional and cryogenic
cooling. Procedia Cirp. 2016. 45:67-70.
[10] Isakson, S., Sadik, M. I., Malakizadi, A., Krajnik, P. Effect of cryogenic
cooling and tool wear on surface integrity of turned Ti-6Al-4V. Procedia
CIRP. 2018. 71:254-259.
[11] Pusavec, F., Lu, T., Courbon, C., Rech, J., Aljancic, U., Kopac, J.,
Jawahir, I. S. Analysis of the influence of nitrogen phase and surface heat
transfer coefficient on cryogenic machining performance. Journal of
materials processing technology. 2016. 233:19-28.
[12] Salame, C., Bejjani, R., Marimuthu, P. A better understanding of
cryogenic machining using CFD and FEM simulation. Procedia CIRP.
2019. 81:1071-1076.
[13] Prevéy, P. S. The use of Pearson VII distribution functions in X-ray
diffraction residual stress measurement. Advances in X-Ray analysis. 1986.
29:103-111.
[14] Ramesh, A., Melkote, S. N. Modeling of white layer formation under
thermally dominant conditions in orthogonal machining of hardened AISI
52100 steel. International Journal of Machine Tools and Manufacture.
2008. 48/3-4:402-414.
[15] Poulachon, G., Moisan, A., Jawahir, I. S. On modelling the influence of
thermo-mechanical behavior in chip formation during hard turning of
100Cr6 bearing steel. CIRP Annals. 2001. 50/1:31-36.
[16] Zorev, N.N. Inter-relationship between shear processes occurring along
tool face and shear plane in metal cutting. International research in
production engineering. 1963. 49:143-152.
[17] Hong, S. Y., Ding, Y., Jeong, W. C. Friction and cutting forces in
cryogenic machining of Ti–6Al–4V. International Journal of Machine
Tools and Manufacture. 2001. 41/15:2271-2285.
[18] Pu, Z., Umbrello, D., Dillon Jr, O. W., Lu, T., Puleo, D. A., Jawahir, I. S.
Finite element modeling of microstructural changes in dry and cryogenic
machining of AZ31B magnesium alloy. Journal of Manufacturing
Processes. 2014. 16/2:335-343.
[19] Bogajo, I. R., Tangpronprasert, P., Virulsri, C., Keeratihattayakorn, S.,
Arrazola, P. J. A novel indirect cryogenic cooling system for improving
surface finish and reducing cutting forces when turning ASTM F-1537
cobalt-chromium alloys. The International Journal of Advanced
Manufacturing Technology. 2020. 111/7:1971-1989.
[20] Tarragó, J. M., Dorvlo, S., Al-Dawery, I., Llanes, L. M. Strength
degradation of cemented carbides due to thermal shock. In Proceedings of
the Euro PM2015 Congress & Exhibition. 2015.
[21] Ortiz-de-Zarate, G., Sela, A., Saez-de-Buruaga, M., Cuesta, M.,
Madariaga, A., Garay, A., Arrazola, P. J. Methodology to establish a hybrid
model for prediction of cutting forces and chip thickness in orthogonal
cutting condition close to broaching. The International Journal of
Advanced Manufacturing Technology. 2019. 101/5-8:1357-1374.
[22] Sela, A., Ortiz-de-Zarate, G., Arrieta, I., Soriano, D., Aristimuño, P.,
Medina-Clavijo, B., Arrazola, P. J. A mechanistic model to predict cutting
force on orthogonal machining of Aluminum 7475-T7351 considering the
edge radius. Procedia CIRP. 2019. 82:32-36.
[23] Kaynak, Y., Tobe, H., Noebe, R. D., Karaca, H. E., Jawahir, I. S. The
effects of machining on the microstructure and transformation behavior of
NiTi Alloy. Scripta Materialia. 2014. 74:60-63.
[24] Pu, Z., Outeiro, J. C., Batista, A. C., Dillon Jr, O. W., Puleo, D. A.,
Jawahir, I. S. Enhanced surface integrity of AZ31B Mg alloy by cryogenic
machining towards improved functional performance of machined
components. International journal of machine tools and manufacture. 2012.
56:17-27.