Высокопроизводительное сухое шлифование кругом с определённым рисунком зерен

CIRP Annals - Manufacturing Technology 57 (2008) 357–362
Contents lists available at ScienceDirect
CIRP Annals - Manufacturing Technology
journal homepage: http://ees.elsevier.com/cirp/default.asp
High-performance dry grinding using a grinding wheel with
a defined grain pattern
J.C. Aurich (2)a,*, P. Herzenstiel a, H. Sudermann a, T. Magg b
a
b
Institute for Manufacturing Technology and Production Systems, University of Kaiserslautern, 67653 Kaiserslautern, Rheinland-Pfalz, Germany
Diamant-Gesellschaft Tesch GmbH, Ludwigsburg, Germany
A R T I C L E I N F O
A B S T R A C T
Keywords:
Grinding
Tool
Dry Machining
This paper presents the potential of a superabrasive electroplated grinding wheel with a defined grain
pattern for dry surface grinding operations. The grinding wheel’s grain pattern is developed by kinematic
simulation with special focus on the undeformed chip thickness. Current experimental investigations of
dry grinding operations of hardened heat-treated steel are carried out with a material removal rate of
Q 0w ¼ 70 mm3 =mm s. The measured grinding forces, workpiece temperatures, as well as workpiece
surface quality and workpiece integrity are compared to wet grinding and a standard superabrasive
electroplated grinding wheel as a reference process.
ß 2008 CIRP.
1. Introduction
Compared to standard machining operations, i.e., milling,
drilling, and turning, grinding processes require the largest amount
of energy for removing a specific material volume. Most of the
energy is transferred into heat caused by internal (material
deformation) and external (contact of grinding wheel and workpiece) friction. Hence, high-performance grinding processes need
high quantities of cooling lubricants to protect the workpiece from
surface damages and to increase the lifetime of the grinding wheel.
Therefore, many research efforts are directed to reduction of
process heat and forces by optimization of the grinding wheel
properties [1–6].
The introduction of dry grinding would save large quantities of
resources (e.g., cooling lubricants, filter units) and energy (e.g., for
pumping and acceleration of the fluid). The dimension of grinding
machines could be reduced due to the absence of cooling lubricant
filters and tanks. However, the main functions of the cooling
lubricant (chip transportation and thermal stabilization of the
machine system) have to be substituted for such machine systems.
Dry grinding processes provide excellent possibilities for optical
measurements, direct cutting force measurement, and the
simplified determination of process heat flows.
The following investigations on dry grinding processes have
been reported [7–11]. External cylindrical dry grinding of 100Cr6V
using a standard electroplated grinding wheel with cutting speeds
up to 300 m/s and specific material removal rates up to 80 mm3/
mm s was carried out by Klocke and Bücker. Material adhesion on
the grinding wheel reduced the lifetime of the wheel and resulted
in surface damage of the workpieces [7]. Voll investigated
cylindrical plunge grinding of C45 using an aluminum oxide
abrasive wheel with a maximum depth of cut of 5 mm performing
* Corresponding author.
0007-8506/$ – see front matter ß 2008 CIRP.
doi:10.1016/j.cirp.2008.03.093
specific material removal rates up to 15 mm3/mm s [8]. Surface
grinding operations of 16MnCr5 and 42CrMo4 using resin- and
ceramic-bonded aluminum oxide abrasive grinding wheels with
cutting speeds up to 45 m/s, depths of cut up to 800 mm and
specific material removal rates up to 3.33 mm3/mm s were
reported by Brinksmeier et al. [9]. Tawakoli et al. conducted
surface grinding operations of 100Cr6 using a structured resinbonded grinding wheel. The surface of the abrasive layer was
decreased to 25%, 50% and 75% by a special dressing process to
reduce the contact area with the workpiece. Cutting speeds up to
40 m/s, depths of cut up to 25 mm, and specific material removal
rates up to 0.625 mm3/mm s were used for the investigations [10].
Due to the relatively small contact area between workpiece and
grinding wheel in external cylindrical grinding the application of
dry grinding with specific material removal rates up to 80 mm3/
mm s could be achieved. In contrast, the large contact area for
surface grinding limited the material removal rate to less than
10 mm3/mm s.
This paper presents the potential of a superabrasive electroplated grinding wheel with a defined grain pattern for dry surface
grinding operations extending the Q 0w limit up to 70 mm3/mm s.
The development of the grinding wheel is based on fundamental
works of Warnecke, Braun and Aurich [1,2].
2. Modeling and simulation
To optimize grinding processes, the complex relationships
between system parameters, machining parameters, process
parameters and work results have to be considered. Therefore,
the relevance of modeling and simulation of grinding processes has
significantly increased in the last years. A general overview about
the numerous models and simulations in grinding is given by
Brinksmeier et al. [12]. In the presented work, a kinematic
simulation originally developed by Warnecke and Zitt [13,14]
and enhanced by Warnecke, Braun and Aurich [1,2] focusing on
358
J.C. Aurich et al. / CIRP Annals - Manufacturing Technology 57 (2008) 357–362
Fig. 1. Grain geometry and distribution.
Fig. 3. Simulation results of numerical dressing.
grinding wheels with a defined grain pattern was applied and
further developed.
Fig. 1 shows the grain geometry and distribution for the
computed model of the grinding wheel. The modeled grain size is
similar to FEPA standard B301. Based on microscopic analysis of
blocky grains used for manufacturing the grinding wheel with a
defined grain pattern, a ratio of 1.13 of the long- and short-grain
axis and a ratio of 3:1 of cubic and tetrahedral basic grain geometry
was modeled. To generate more realistic grain shapes, planes with
randomized orientation are cut off the grain’s basic geometries
[14].
The design and the notation of the grain pattern are depicted in
Fig. 2. Based on the applied manufacturing process, a grain
protrusion height of 35% of the grain size is taken for modeling the
grinding wheel. For an average short grain axis length of 305 mm
and a standard deviation of 30 mm the average grain protrusion
height is 107 mm.
Within the simulation, grinding wheel wear is not considered.
Hence, the workpiece material is removed by grains with high
grain protrusion even if the chip thickness is more than 10 mm or
15 mm. However, for real grinding processes grains with very high
chip thickness wear out very rapidly. The number of kinematic
grains is increasing and the material removal process respectively
the grinding forces are distributed to many grains until a stationary
grinding process is reached. To simulate the stationary grinding
process, a numerical dressing process of the modeled grinding
wheel was performed. In detail, the grain protrusion height of
every single grain is ideally cut down to a specified limit. Fig. 3
shows the impact of different dressing heights upon the
percentage of kinematic grains out of the total number of grains
during a stationary grinding process with constant setup parameters compared to the undressed grinding wheel.
These results are used for the dressing process of the
manufactured prototype and enable grinding with many kinematic
grains starting from the first workpiece. Hence, a more uniform
stress on the single grains can be achieved resulting in a steady
state grinding process and a reduction of grain breakouts. In
general, grain breakouts increase the stress to following grains.
Hereby, the lifetime of a grinding wheel with a reduced number of
grains is much more affected than of a standard electroplated
Fig. 2. Notation of modeled grain patterns [2].
Fig. 4. Modeled grain pattern and simulation results.
grinding wheel. To avoid a chain reaction of grain breakouts on a
circumferential line on the grinding wheel with a defined grain
pattern, as observed before by Braun et al. [2], the number of grains
in each line was doubled.
Fig. 4 shows the modeled grain pattern numerically dressed to a
maximum grain protrusion height of 110 mm. In comparison to the
Fig. 5. Reference wheel and prototype.
J.C. Aurich et al. / CIRP Annals - Manufacturing Technology 57 (2008) 357–362
Table 1
Specification of grain pattern
Grain
Grain
Grain
Grain
Grain
size (DIN ISO 6106)
line angle (a)
line distance (Dx)
distance (Dz)
line displacement (Dzv)
359
4. Experiments
Blocky B301
608
1500 mm
300 mm
100 mm
real grinding wheel (Fig. 5) the dimensions of the modeled grinding
wheel are 400 mm 5 mm instead of 400 mm 12 mm to save
computing time.
The simulated process is a stationary surface grinding process
of one grinding wheel revolution with Q 0w ¼ 70 mm3 =mm s and
depths of cut of 0.75 mm up to 5 mm. Concerning the simulation
results, the specific number of momentarily active grains (N0mom )
and the mean value of the average chip thickness (hcu,med) of the
kinematic grains during one simulated grinding wheel revolution
are displayed. N0mom increases with increasing depth of cut of
N0mom ¼ 19 mm1 for ae = 0.75 mm up to N 0mom ¼ 36 mm1 for
ae = 5 mm. Due to the constant Q 0w the removed workpiece
material volume within a time interval is constant and therefore
independent of the depth of cut. Hence, the material removal
process is distributed to more grains with increasing depth of cut
which explains the decrease of the simulated average chip
thickness. The average chip thickness falls from hcu,med = 0.156 mm
for ae = 0.75 mm to hcu,med = 0.09 mm for ae = 5 mm.
3. Manufacturing and preparation of the prototype
Based on the results of the simulation, the grain pattern for a
prototype grinding wheel with the dimensions 400 mm 12 mm
was specified (Table 1).
To create the grain pattern, masking technologies were used for
preliminary adhesion of the grains on the steel wheel hub. For
completion, a standard electroplating process was applied. About
35,000 grains were placed to form the abrasive layer of the wheel
(Fig. 5). Due to the prototypical manufacturing technology, up to
five grain lines at three positions are missing on the wheel (less
than 2% of the total number of grain lines). However, the grinding
wheel could be used for the whole series of experiments without
further loss of grain lines.
By using masking technology for manufacturing the presented
prototype significant tool cost reduction could be achieved in
comparison to the initially developed prototype wheel by
Warnecke, Braun and Aurich [1,2], where each single grain was
positioned manually. However, the tool costs of the prototype are
still considerably higher in comparison to the used reference tool.
After mounting the grinding wheel to the machine spindle
(radial runout less than 1.5 mm) a touch-dressing operation was
carried out. A rotating diamond dresser of 120 mm in diameter
was used at a dressing speed ratio of qd = 0.6. The axial feed rate
was 10 mm/rev and the infeed increment 1 mm. The uniformity of
the maximum grain protrusion height as a result of the dressing
process was checked by a roughness measurement in the negative
profile of the grinding wheel. The negative profile was prepared by
a plunge cut grinding process into a steel sheet of 1 mm thickness.
The average roughness height (Ra) and the 10-point height (Rz) of
the steel sheet were measured. Before dressing the roughness was
Ra = 10.59 mm and Rz = 52.29 mm and after dressing Ra = 0.44 mm
and Rz = 3.73 mm. The average grain protrusion height of the
grinding wheel was measured by analysis of 30 single grains of the
grinding wheel’s surface replica using confocal microscopy. The
replicas were prepared at specified positions on the grinding
wheel. The average grain protrusion height was 118 mm before
and 102 mm after dressing. According to the simulation results of
Fig. 3, the number of kinematic grains increases to approximately
60% of the total number of grains due to the dressing process.
Hence, the grinding forces are distributed uniformly to the grains
and a stationary wear behavior of the grinding wheel can be
assumed.
The surface up grinding experiments were carried out on a high
efficiency grinding machine with high stiffness. In addition to the
grinding wheel with a defined grain pattern (prototype), a standard
superabrasive electroplated grinding wheel of the same geometry
and grain size was used as reference wheel (Fig. 5).
As workpiece material heat-treated steel AISI 4140H
(42CrMo4V, DIN EN 10132-3) of 60 2 HRC with the dimensions
125 mm 50 mm 10 mm was used. Cooling lubricant during wet
grinding was grinding oil.
The grinding process and its results were monitored by the
following measurement techniques: grinding forces (normal and
tangential direction) by dynamometer, effective spindle power by
non-contact hall effect probe, workpiece temperature by thermocouple sensor, and surface roughness by tactile measurements.
Within the experimental investigations three different setups
with Q 0w ¼ 70 mm3 =mm s and varying depth of cut ae = 1, 2, 3 and
4 mm with constant cutting speed of vc = 100 m/s were performed
(Table 2).
4.1. Grinding forces, spindle power and workpiece temperatures
In Fig. 6a, the comparison of the net grinding forces and the net
spindle power is depicted, i.e., impact of the cooling lubricant for
wet grinding has been taken out. Therefore, in separate grinding
passes of each parameter combination for wet grinding the feed
motion during stationary grinding was stopped and grinding forces
as well as spindle power caused by the cooling lubricant were
recorded. For dry grinding the idling power was deducted. The
displayed values of force and power are average values of at least
three measurements during the stationary grinding process and
are low-pass filtered (20 Hz). The maximum standard deviation of
the normal forces is 25 N, tangential force 15 N and spindle power
0.5 kW. The workpiece temperatures during stationary grinding
were measured by thermocouple sensors (diameter 0.5 mm)
laterally fixed in 0.6 mm holes at a distance of 0.5 mm to the
ground surface in the middle of the workpiece in the grinding
direction (Fig. 6b). Distance variations of the thermocouple sensor
and the ground surface were normalized by interpolating the
distance–temperature curve of the corresponding workpiece.
For wet grinding using the reference wheel and the prototype,
an increase of grinding forces and power with increasing depth of
cut can be observed in Fig. 6a. The prototype (Table 2, setup 2)
shows lower grinding forces and spindle power in comparison to
the reference wheel (Table 2, setup 1). Concerning the normal and
tangential grinding forces, a reduction of 30–40% and 20–35%,
respectively could be achieved. The reduction of spindle power
was around 16–27%. For dry grinding the forces and the spindle
power are nearly constant with a further reduction in comparison
to wet grinding.
Concerning workpiece temperatures (Fig. 6b) during dry
grinding, a significant temperature increase from 530–1050 8C
can be observed for increasing depth of cut. Wet grinding processes
show nearly constant workpiece temperatures, apart from a minor
deviation for ae = 2 mm. For wet grinding processes using the
prototype wheel, lower workpiece temperatures in comparison to
the reference wheel occur.
The results for wet grinding consistently show a better process
behavior for the prototype. Due to an increased ratio of material
removing grains to ploughing grains for the prototype, the cutting
Table 2
Setups of grinding experiments
Setup
Grinding wheel
Lubricant use
1
Standard superabrasive grinding wheel
(reference wheel)
Defined grain pattern (prototype)
Defined grain pattern (prototype)
Yes
2
3
Yes
No
360
J.C. Aurich et al. / CIRP Annals - Manufacturing Technology 57 (2008) 357–362
Fig. 7. Material adhesion on grinding wheel.
due to lower feed rates were expected. Nevertheless, the workpiece temperatures remained nearly constant. This can be
explained by higher heat removal by the cooling lubricant. The
absence of cooling lubricant during dry grinding led to a higher
heat flux into the workpiece and increased workpiece temperatures in comparison to wet grinding.
While workpiece temperatures in wet grinding are constant for
increasing depth of cut, the temperatures in dry grinding increase.
The reason is an increased heat flux into the workpiece due to the
increasing contact length combined with the decreasing feed rate.
The higher heat flux heats a bigger volume of workpiece material
and the grindability of the thermally softened material is
increased. Therefore, grinding forces and spindle power in dry
grinding do not increase.
As a side effect of higher temperatures, adhesion of workpiece
material on the prototype wheel occurred with high depths of cut
(Fig. 7). Thus, the friction between grinding wheel and workpiece
further increased and contributed to higher temperatures (Fig. 6b).
However, the material adhesion did not damage the prototype and
even after many grinding operations without active cleaning of the
grinding wheel’s surface the material adhesion did not accumulate.
Instead, a stationary self-cleaning process occurred. Existent
material adhesion completely disappeared after grinding operations with low depth of cut of ae = 1 mm or 2 mm.
After extensive experimental research a specific material
removal of V 0w ¼ 80; 000 mm3 =mm was carried out with the
prototype, including V 0w ¼ 45; 000 mm3 =mm of dry grinding
operations. Concerning wheel wear only a few grain breakouts
are detected by scanning electron microscopy (SEM) analysis.
Primarily, abrasive wear of the grains occurred.
4.2. Workpiece surface quality
The workpiece surface roughness was measured orthogonal
to the grinding direction in the middle of the ground workpiece
surface. Fig. 8 shows the mean value of three measurements of
the 10-point height Rz. The maximum standard deviation is less
than 1 mm.
Fig. 6. Experimental results. (a) Grinding forces and spindle power and (b)
workpiece temperatures.
process is improved. Therefore, process forces and power are
reduced. The reduced amount of ploughing grains results in a
lower generation of heat and consequently lower workpiece
temperatures. Furthermore, larger chip space enables a better chip
transportation, a better lubrication of the contact zone, and
therefore a higher heat removal rate.
The increase of grinding forces and power with a larger depth of
cut for wet grinding is based on increased contact lengths leading
to more active grains and consequently increased friction. Initially,
increased heat generation and higher heat flux into the workpiece
Fig. 8. Workpiece surface roughness.
J.C. Aurich et al. / CIRP Annals - Manufacturing Technology 57 (2008) 357–362
The growing surface roughness with increasing depth of cut for
dry grinding and the nearly constant surface roughness for both
wet grinding processes resemble a good correlation to the
workpiece temperature in Fig. 6b. Due to higher workpiece
temperatures and a softer workpiece material the cutting process
at microscopic level is increasingly substituted by ploughing and is
therefore less efficient. The best surface quality is achieved by the
application of the reference wheel caused by higher amounts of
grains with low chip thickness. The intensified material adhesion
for higher temperatures in dry grinding contributes to rougher
workpiece surfaces. Visual grinding burn did not occur either for
wet or for dry grinding experiments using the prototype wheel.
However, the surface of the dry grinding experiments compared to
wet grinding was less shiny. Apparently, higher surface temperatures and the missing cooling lubricant enhanced the oxidation
process on the surface.
4.3. Workpiece surface integrity
For investigation of the workpiece surface integrities metallographic sections of the workpieces were prepared for microhardness and microstructure analysis. For the microstructure
analysis the sections were etched for 20 s using 3% of alc. HNO3.
The material of the workpiece AISI 4140H (42CrMo4V, DIN EN
10132-3) austenitizes at around 800 8C and tempers at around
550 8C. High cooling rates of austenitized material result in a
hardness increase by martensitic hardening of the material.
Tempering leads to a hardness decrease of the material. Hence,
three different zones (Fig. 9) can occur on the metallographic
sections dependant on the workpiece temperature and the cooling
rate: rehardened zone, tempered zone and unmodified workpiece
core.
The different zones are identified by micro-hardness analysis
(Fig. 9a) and their different microstructures verified by light
microscope and SEM images. The transition of the rehardened zone
to the tempered zone is identified by an abrupt hardness decrease
to around 400 HV which allows for a very accurate measurement of
the thickness of the rehardened zone. The transition of the
tempered zone to the workpiece core cannot be identified clearly
due to the continuous hardness increase towards the workpiece
core. Therefore, a hardness state of 600 HV, which is almost the
hardness of the workpiece core, is defined as transition to the core.
In Fig. 9a, the micro-hardness of a workpiece with rehardened
zone – exemplarily shown for dry grinding with ae = 2 mm – is
depicted. The maximum measured micro-hardness of the rehardened layer is 800 HV. The hardness of the tempered zone starts
around 380–400 HV, directly below the rehardened zone, and
increases towards the workpiece core to 650 HV.
For dry grinding operations, the rehardened and tempered zone
show maximum thickness in comparison to the wet grinding
processes (Fig. 9b). The rehardened and the tempered zone
significantly increase from 400 mm to 1250 mm, respectively from
1000 mm to 6800 mm with increasing depth of cut. Although no
cooling lubricant is used, fast cooling rates caused by a high heat
flux into the workpiece and air cooling, enable martensitic
hardening. For wet grinding using the prototype, the workpieces
show no rehardened zones. The tempered zones increase from
1000 mm up to 1900 mm with increasing depth of cut. Wet
grinding using the reference wheel results in rehardened zones of
around 80 mm thickness for ae = 1 mm and 2 mm. For ae = 3 mm
and 4 mm no rehardened zones occured. The tempered zone
increases with higher depths of cut starting from 1000 mm up to
3000 mm.
The thickness of the identified rehardened and tempered zones
increases with higher heat flux into the workpiece as a
consequence of higher workpiece temperatures. Hence, a good
correlation to the measured workpiece temperatures occurred.
Selected X-ray residual stress measurements on the workpiece
surface (maximum penetration depth 5.5 mm) using Cr Ka
radiation and sin2 C method for stress calculation are displayed
Fig. 9. Microstructure and micro-hardness analysis.
Fig. 10. Residual stresses on the workpiece surface.
361
362
J.C. Aurich et al. / CIRP Annals - Manufacturing Technology 57 (2008) 357–362
in Fig. 10. Residual stresses in grinding direction (sII) and normal to
grinding direction (s?) were measured. The generated residual
stresses reveal only compressive stresses at the workpiece surface.
In grinding the residual stress state depends on the superposition of thermal and mechanical influences, accounting for
tensile and compressive residual stresses, and material influences,
e.g., phase transformations leading to tensile (volume decrease) or
compressive stresses (volume increase) [11].
For dry grinding the occurring martensitic phase transformation (volume increase) within the rehardened zone and the
mechanical influences overcompensate the thermal-induced
tensile stresses and lead to compressive residual stresses. The
same reasons are responsible for the occurrence of compressive
stresses during wet grinding for ae = 1 mm using the reference
wheel where a rehardened zone was also identified.
For wet grinding using the prototype (ae = 1 mm and 4 mm) and
the reference wheel (ae = 4 mm) instead of rehardened zones
tempered zones occurred. Tempered zones are generated by
martensitic to ferrite/perlite phase transformations causing a
volume decrease and therefore accounting for tensile stresses.
Since compressive residual stresses were measured on the
workpiece surface the mechanical influences of the wheel–
workpiece contact predominated the residual stress formation.
5. Conclusion and outlook
Based on kinematic simulation, a prototype grinding wheel
with a defined grain pattern was developed and high-performance
dry surface grinding operations were carried out. To prove the
capability of the prototype wheel, wet grinding operations of the
prototype were compared to a standard electroplated superabrasive grinding wheel. The prototype showed better process
behavior with lower grinding forces and power of up to 40% and
lower workpiece temperatures in wet grinding. Grinding power,
forces and workpiece temperatures as well as workpiece surface
roughness, microstructure, micro-hardness and residual stresses
were measured. The application of the prototype wheel for dry
grinding led to higher workpiece temperatures and a higher
alteration of surface layers in comparison to wet grinding. Higher
workpiece temperatures improved the grindability of the material
and decreased process forces and power. However, the workpiece
quality of wet grinding process could not be reached. Finally,
extensive experimental research of wet and dry surface grinding
operations with Q 0w ¼ 70 mm3 =mm s and V 0w ¼ 80; 000 mm3 =mm
did not damage the prototype and showed only minor abrasive
wear at the grinding wheel.
Future investigations will include the measurement of residual
stress depth profiles for a better understanding of the alteration of
surface layer. Due to high thermal impacts during dry grinding part
form errors, e.g., workpiece flatness, will be measured. Moreover,
dry grinding of different materials, e.g., high-temperature alloys
will follow.
Acknowledgements
This research and development project was partly funded by
the German Federal Ministry of Education and Research (BMBF)
within the Framework Concept ‘‘Research for Tomorrow’s Production’’ and managed by the Project Management Agency Forschungszentrum Karlsruhe, Production and Manufacturing
Technologies Division (PTKA-PFT).
References
[1] Aurich JC, Braun O, Warnecke G (2003) Development of a Superabrasive
Grinding Wheel With Defined Grain Structure Using Kinematic Simulation.
Annals of the CIRP 52(1):275–280.
[2] Braun O, Warnecke G, Aurich JC (2005) Simulation-based Development of a
Superabrasive Grinding Wheel with Defined Grain Structure. Proceedings
NAMRC North American Manufacturing Research Conference, vol. 33, 351–
358.
[3] Pinto FW, Wegener K, Kuster F, Wunder S, Kiser M (2005) Model to Optimise
the Grain Pattern of Grinding Discs with Controlled Positions of the Abrasives.
Proceedings of the 8th CIRP International Workshop on Modeling of Machining
Operations, Chemnitz, Germany, 349–359.
[4] Tawakoli T (2001) Innovatives System zum Schleifen, Abrichten und Fräsen.
IDR 35:48–54.
[5] Abele E (2007) Mehrtechnologie orientierte rekonfigurierbare Werkzeugmaschine, Final report METEOR 2010, BMBF.
[6] Chattopadhyay AK, Hintermann HE (1994) On Performance of Brazed SingleLayer CBN Wheel. Annals of the CIRP 43(1):275–280.
[7] Klocke F, Bücker C (1996) Trockenschleifen-Grenzbetrachtung zur Kühlschmierstoffreduzierung. IDR 1(96):26–35.
[8] Voll M (1999) Schleifprozesse mit Korundwerkzeugen für die Trockenbearbeitung
auslegen, vol. 105 (31). Maschinenmarkt, Würzburg. pp. 24–27.
[9] Brinksmeier E, Brockhoff T, Walter A (1997) Minimalmengenschmierung und
Trockenbearbeitung beim Schleifen, vol. 52 (3). Härterei-Technische Mitteilungen-HTM. pp. 166–170.
[10] Tawakoli T, Westkaemper E, Rabiey M (2006) Dry Grinding by Special Conditioning. Proceedings CIRP International Conference on High Performance Cutting, Vancouver, .
[11] Brockhoff T (1999) Grind-Hardening: A Comprehensive Overview. Annals of
the CIRP 48(1):255–260.
[12] Brinksmeier E, Aurich JC, Govekar E, Heinzel C, Hoffmeister H-W, Peters J,
Rentsch R, Stephenson DJ, Uhlmann E, Weinert K, Wittmann M (2006)
Advances in Modeling and Simulation of Grinding Processes. Annals of the
CIRP 55(2):667–696.
[13] Warnecke G, Zitt U (1998) Kinematic Simulation for Analyzing and Predicting
High-Performance Grinding Processes. Annals of the CIRP 47(1):265–270.
[14] Zitt U-R (1999) Modellierung und Simulation von Hochleistungsschleifprozessen,
Dissertation, University of Kaiserslautern.