Journal Of Materials Processing Technology 202 (2008)

j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 2 ( 2 0 0 8 ) 328–338

journal homepage: www.elsevier.com/locate/jmatprotec

Accuracy of hard turning
´ a , B. Karpuschewski b , K. Gyani a , V. Bana c,∗
J. Kundrak
a
b
c

Department of Production Technology, University of Miskolc, Miskolc, Hungary
Institute for Production Technology and Quality Assurance, Otto-von-Guericke University, Magdeburg, Germany
Advanced Technology Centre (ATC), Philips DAP bv., Drachten, The Netherlands

a r t i c l e

i n f o

a b s t r a c t

Article history:

Nowadays, hard turning is frequently used to replace grinding. The economic benefits of

Received 16 October 2006

hard turning are obvious but for achievable accuracy the situation is somewhat ambiguous.

Received in revised form

Although machine tool factories offer lathes with the same accuracy as grinding machines

28 August 2007

in some cases problems may arise in keeping the prescribed geometrical accuracy. Inves-

Accepted 3 September 2007

tigations were performed in a working environment in order to determine the attainable
size, form and positional accuracy obtained with hard turning. Error sources of machining
errors that occurred in hard turning and in grinding were taken into account, giving typical

Keywords:

differences between the two processes. In the parts produced in series, size deviations were

Hard turning

measured as well as out-of-roundness, cylindricity error and parallelism error of the bore’s

Accuracy

generatrices. The workpieces used for the investigation are disc-type parts with bores, i.e.,

Form error

gears that are built into transmissions. Our first measuring series evaluates the achievable

Size error

accuracy with hard turning while the second includes the comparison of grinding with hard
turning. The most important error sources are identified. We present measures for keeping
prescribed tolerances and propose methods for eliminating the means error source.
© 2007 Elsevier B.V. All rights reserved.

1.

Introduction

The primary task of hard turning – as a finishing operation –
is to ensure the quality and reliability of the parts. The quality
criteria of hard turning as a cutting operation can be found
in the technical drawings. The most important quantities are
geometrical accuracy, surface topography, and the integrity of
the subsurface layer.
Geometrical accuracy includes size errors, form errors and
positional errors. Surface topography covers the drawing,
the roughness and the bearing curve of the surface. Surface
integrity describes changes in the physical properties of the
material that result from machining (Barry and Byrne, 2002).
For analysis of the geometrical accuracy, four typical characteristics of hard turning – as opposed to grinding – must be
outlined.

∗

Corresponding author.
E-mail address: (V. Bana).
0924-0136/$ – see front matter © 2007 Elsevier B.V. All rights reserved.
doi:10.1016/j.jmatprotec.2007.09.056

These are: (1) the significantly higher cutting force, (2) the
omission of coolant, (3) the single point form generation, and
(4) the minimum value of the depth of cut.
The cutting force occurring in hard turning is higher than
that in conventional turning or grinding. The passive force
occurring in hard turning – the component perpendicular to
the cutting speed – is a multiple of the main cutting force,
while in traditional turning it is only a fraction of this value.
The extraordinarily high passive force, which contributes to
the material removal, significantly loads the elements of the
machining system, causing elastic deformation and deteriorating the machining accuracy. The disadvantageous effect of
the high passive force must be compensated for by an increase
in machine tool rigidity.
Hard turning can be done in dry conditions at relatively
high speed. The relative high friction coefficient and the pas-


j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 2 ( 2 0 0 8 ) 328–338

sive force cause a significant friction force which transforms
into heat. The other source of the generated heat is the high
cutting speed. This cannot be reduced, because PcBN is only
effective at high temperatures. The high temperature generated during material removal causes thermal expansion of the
workpiece, which also deteriorates the machining accuracy.
The disadvantageous effect of the hot swarf falling on the elements of the machine tool must be compensated for with the
increase of “heat rigidity” of the machine (Schmidt and Dyck,
2000).
The surface generating element of hard turning is the
single-point tool tip, which shapes the surface of the workpiece and is accompanied by significant force and heat
effects. Under such conditions the single-point tool tip reacts

sensitively to any irregularities. It abreacts the allowance
distinctions, the hardness differences and the other heterogeneities of the material with the creation of machining errors.
On the other hand, grinding, where the surface generating
element is the linear generatrix of the wheel and the forces
are also lower, hardly detects any inhomogeneities in its path;
it simply eliminates them and creates a more accurate surface.
The fourth specification influencing the accuracy is the
depth of cut. In hard turning this cannot be reduced arbitrarily,
although this is possible in grinding. Moreover, grinding can
also be performed with zero depth of cut, called spark-out, to
eliminate any deflections of the system by continuously reducing the forces. Because of the necessity of a minimum depth of
cut, hard turning is followed by higher forces than in grinding,
even in the finest smoothing operations.
Due to the four typical characteristics of hard turning mentioned above, the tool tip is exposed to a substantially more
intensive physical–mechanical load than a single grain of the
grinding wheel. The density of energy transmitted into the
workpiece on the tool tip is much higher. Therefore, the stationary state is more difficult to maintain and the accidental
error sources arising in the material removal process are more
difficult to handle than in grinding. Most unfavourable effects
can be eliminated by an increase in the robustness of the
machining system. The main element of the machining system is the machine tool, which must possess extraordinarily
high static and dynamic stiffness and must also withstand
the heating effect of the hot swarf falling upon it. A second element is the clamping device for the workpiece. It has
a large influence on the geometrical accuracy in disc-type
components, which sometimes possess thin walls. Extremely
rigid clamping devices and deliberate clamping forces systems must be applied. The third element is the tool-clamping
device together with the tool, which offers few possibilities to
increase the rigidity. The fourth element of the system is the
workpiece, which is given and bears the errors issuing from
¨

the motions of system elements (Tonshoff
et al., 1997).
Unintended error sources occurring in hard turning can be
divided into two classes: error sources depending on the load
and error sources independent of the load.
Error sources depending on the load:

1 Cutting force, which creates elastic deformations;
2 Cutting heat, which causes distortions;

329

3 Tool wear, which increases the force and the heat;
4 Insufficient static rigidity of the machining system;
(a) The machine tool with insufficient stiffness becomes
deformed from the force.
(b) The workpiece’s clamping device of inadequate rigidity
is deformed.
(c) The weak tool holder and tool are bent.
(d) A workpiece with unsatisfactory stiffness is deformed
by the clamping force.
5 Because of the inadequate dynamic stiffness of the machining system, low-frequency oscillations may cause form
error.
Error sources independent the load:
1 Uneven allowances cause force fluctuation and form error
and release stresses.
2 Inhomogeneities of the workpiece material cause from
error by the force fluctuation.
3 Manner of the surface generation: in turning a point, in
grinding a line generates the surface.

4 Construction of the clamping device and the deforming
effect of the clamping force may be a significant source of
form errors. Instead of concentrated force it is more suitable
to clamp with distributed force.
5 Number of clampings: hard turning has an advantage
because with the single point tool and one clamping various surfaces can be machined. In grinding usually several
clampings are necessary, which is a source of significant
form and positional errors.

2.

Experimental conditions

2.1.
Type and main sizes of the components and
accuracy prescriptions
This investigation was performed in a working environment
on gears with bores machined in batch. The gears are built
into the transmissions of motor vehicles, and are disc-type
components of different sizes. The material of the gears is
case-hardened steel: 16MnCr5 (AISI 5115) with hardness 62 ± 2
HRC. Four out of 20 measured gears are presented with the
measuring setup in Fig. 1 in order to show the sizes and types
investigated.
The accuracy prescription for the gears applies to the size,
form and positional accuracy. In many cases the size accuracy
of the bores is IT6, occasionally IT5 or IT7. For form errors,
the out-of-roundness of the bore and the flatness error of the
faces are prescribed. Furthermore, the parallelism of the bore’s
generatrices and the axial run-out of the faces has to be kept

within certain limits. The specifications of form and positional
accuracy are set by the internal standards of the factory on
the basis of functional conditions. These are much stricter
than the prescriptions in the general standards. For instance,
to the Hungarian standard MSZ ISO 2768-2:1991 permits outof-roundness equal to half of the diameter, which is 2–3 ␮m
larger than the value prescribed by the factory.
Table 1 summarizes the prescriptions of accuracy for hard
turned and ground surfaces for the 10 gears chosen as exam-


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Fig. 1 – Main types and sizes of gears A, B, C: Measuring planes of out-of-roundness.

Table 1 – Prescription of accuracy for hard turned/ground surfaces


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331

Fig. 2 – Interpretation of measuring results (a) Out-of-roundness, (b) cylindricity, (c) parallelism of the bore’s generatrix, (d)
measuring setup for determination of flatness and axial run-out, (e) flatness and (f) axial run-out.

ples. The root diameter/bore diameter (droot /d1 ) and bore
length/bore diameter (L/d1 ) ratios are also shown, since these
play a very significant role in the formation of the measuring
results.

Although the prescriptions of accuracy according to the
internal standards of the factory do not include cylindricity,
cylindricity error was also measured and the alteration of its
value is presented. In positive cases the cylindricity error mirrors the parallelism of the generatrices. As the investigation
of parallelism is done in two-dimensional planes, the result
does not provide any reliable information about the cylindricity error.
The finish hard turning was performed on an advanced
machine tool suitable for the requirements of hard turning and was done in one clamping, in a hydraulic three-jaw
chuck. The special jaws centralize on the pitch circle of

Fig. 3 – Out-of-roundness, cylindricity and parallelism in
hard turning.


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Fig. 4 – Generation of out-of-roundness in hard turning.

the gears and clamp with concentrated force. The clamping
force is not known; only the pressure of the system, which
is 12 bar. The finish grinding was done on several grinding
machines, with several clampings; a three-jaw chuck was
used.

2.2.

Machining conditions

The experimental conditions for machining of gear bores are:
(a)Hard turning
Machine tool:
PITTLER PVSL-2 lathe
Cutting tool:
PcBN CNGA 120408 BNC80
◦
◦
◦
n = −6 , ˛n = 6 , Är = 95 ,
εr = 80◦ , r␧ = 0.8 mm
(CAPTO C5-PCLNL-17090-12)
Technological data:
vc = 120–228 m/min,
f = 0.08–0.1 mm/rev,
ap = 0.1 mm
(b)Grinding
Machine tool:
SI-4/A internal grinder (VEB
Berliner
Werkzeugmachinenfabrik)
Grinding wheels:
CBN wheel 50 × 32 × 20 9A 80
K7 V22 Bay state
CBN wheel 60 × 36 × 20 9A 80
K7 V22 Bay state

Fig. 6 – Size accuracy of hard turned bores machined in
sequence.

Fig. 5 – Generation of cylindricity error in hard turning.


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Technological data:
Wheel speed:
Speed ratio:
Workpiece speed:
Depth of cut:

vc = 25 m/s
q = 80
vw = 18 m/min
ae = 0.020 mm/double stroke,
roughing
ae = 0.020 mm/double stroke,
finishing

Sparking out:
Traverse speed:

Feed rate:
Coolant:

Fig. 7 – Diagrams of hard turned and ground profiles.

333

8–6 double stroke

vf,L = 2200 mm/min, roughing
(traverse speed)
vf,L = 2000 mm/min, finishing
f = 13.7 mm/workrev, roughing
f = 12.0 mm/workrev, finishing
half synthetic, Q = 50 l min−1


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Grinding of the gear bores presented in Table 1 was
performed with internal traverse grinding. The grinding
allowance is 0.4 mm in radial direction. For each piece the
dressing of the wheel is done four times with natural diamond
(0.5 carat). The depth of cut for dressing is ad = 0.01 mm.

3.

Measuring methods and devices

The interpretation of the measuring data and review of the
measuring methods can be seen in Fig. 2. The determination
of the real geometrical forms was done by a 4 mm diameter
stylus head. Out-of-roundness is either defined by the longest
distance between the fitting circle and the given points of the
real profile, or the instrument determines a so-called reference circle and out-of-roundness is determined as a value of
the largest positive local roundness deviation added to the
absolute value of the largest negative local roundness deviation (Fig. 2a). The last method is more frequently used and

it was applied in this study. The reference circle is defined
and written out by the measuring computer on the basis of
the least square sum of the deviations. Cylindricity error is
given as value of the largest positive local cylindricity deviation added to the absolute value of the largest negative
local cylindricity deviation (Fig. 2b). The computer generates
the real cylinder using the circumferential section method:
the instrument determines the out-of-roundness diagrams
in three or four planes perpendicular to the axis and from
them it applies a cylinder superficies built up from straight
lines. This is the real cylinder form. The parallelism error
of the generatrices is the absolute difference in local diameter at the top and the local diameter at the bottom of the
cylindrical future of the two associated lines fitted through
the two generatrix profiles obtained from an intersection of
a plane through the axis of the least squares reference cylinder and the cylindrical feature within the full extent of the
feature (Fig. 2c). The flatness error of the face and its run-out
is determined during one scanning by analysing the writtenout profile in two different manners (Fig. 2d). As the face can
be a plane either with or without run-out, both errors may
appear together. In the measurement of flatness firstly a reference plane must be determined in the profile outstretched
into a plane. The trace line of the reference plane (MMZRP)
is also defined by the least square sum method, by means of
regression analysis (Fig. 2e). After this two reference planes
(OMZRP, IMZRP), parallel to the reference plane are fitted on
the real surface. The distance between these two planes determines the flatness error (FLTt). In the measurement of the
axial run-out of the face the real surface is also touched
with two planes that are perpendicular to the datum axis
rather than parallel with the reference plane (Fig. 2f). The
distance between these two planes defines the axial run-out
(R).
The diameter of the bores was measured by the Derby
Etalon 454 coordinate measuring machine. The applied stylus was the Renishaw Brown Shape TP-ES. This measuring

machine measures the diameter of one circle with three
touches. We repeated it on four circles and their average value
formed the real value for the diameter.

The measurement of out-of-roundness, parallelism, cylindricity, flatness and axial run-out was done with a Talyround
252e roundness measuring machine, as was the preparation
of the diagrams presented in this article.
In the measurement of geometrical accuracy, electronic
filters are important in order to ensure the range of
suitable frequency transmissions, i.e., the number of undulations being set regularly on the circumference, which
can be observed by the instrument during one revolution. We used a Gauss-type filter with characteristic of
1–500 undulations/revolution. This filter senses undulations
up to 500, which are set regularly on the circumference. Compared with Gauss-type filters of 1–50 u/r or 1–15 u/r, this filter
can give a more-detailed idea of the form of the out-ofroundness. It is allowed to compare the gained diagrams only
if they were measured with the same filter.

4.
Out-of-roundness, cylindricity error and
parallelism error in hard turning
The geometrical accuracy of hard turning was investigated in
series production. One series each of three gears was measured; the series contained 285, 60, and 200 pieces. The gears
differ in droot /d1 ratio. Out-of-roundness, cylindricity and parallelism measurements were performed at regular intervals:
after each 25th piece for the 285-piece series, each 5th for the
60-piece series, and each 10th for the 200-piece series. Average
values of these measurements are given in Fig. 3. The measuring results truly reflect the geometrical errors issuing from the
deformation of the workpiece. The “rigidity” of the workpiece,
the weakest element of the system, is defined by the droot /d1
ratio. With the decrease of the droot /d1 ratio, out-of-roundness
and cylindricity error increase. The form errors increase not
only in absolute meaning but also its scattering significantly

rises. However, the parallelism error of the generatrices hardly
changes, nor does its scatter; each measured value remains
much lower than the prescribed 0.006 mm.
Out-of-roundness and cylindricity error are presented in
Figs. 4 and 5, where every measuring result is shown. For the
gear with droot /d1 = 2.83 ratio, which can be regarded as quite
rigid, the out-of-roundness of 0.006 mm prescribed for quality IT6 can be kept, but this cannot be guaranteed for the

Fig. 8 – Summary of measured results.


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gear of droot /d1 = 1.34 ratio. The reason for this is obviously
the clamping chuck, because the clamping method and force
are not suitable. It is mechanical clamping with high pressure instead of magnetic clamping. On the newly developed
lathes, in addition besides 12 bar, two lower pressures are also

335

set with 8 and 4 bar. As can be seen in Fig. 5, the generation
of the cylindricity errors also proves that the wall thickness
(droot /d1 ) must be taken into account in the selection of the
clamping force. Although there are no prescriptions for cylindricity in the drawings, their magnitude must be controlled

Fig. 9 – Diagrams of hard-turned and ground profiles.


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because they significantly influence the function of the gears.
As can be observed in Fig. 5, for the ratio of 2.83 the cylindricity
error does not exceed 0.004 mm but for the ratio of 1.34 it is
3–4 times higher.

5.

Size accuracy of hard turning

The size accuracy of hard turning was measured for three
gears produced in series. The measuring series consisted of
339, 297 and 102 pieces. The result of the measurements is presented in Fig. 6.The magnitude of scattering does not exceed
the whole tolerance range for the bores with accuracy IT6.
This means that on the lathe the accuracy IT5 can be ensured
easily. Moreover, if the possibility of size correction is used
maximally, it is possible to meet more restrictive tolerances
as well, with proper rigidity and good maintenance of the
machine tool. Furthermore, nowadays machine tool factories
produce lathes not only for disc-type parts but also for external, cylindrical, conical and shaped surfaces. Their guaranteed
machining accuracy is IT5 with out-of-roundness of 1 ␮m and
a cylindricity deviation of 2 ␮m.

6.
Formation of accuracy in comparison
with grinding
Many factories have already changed to hard turning and,
therefore, in some cases grinding is only applied as a necessary solution. However, we were able to measure gears
produced with both hard turning and grinding. Fig. 7 introduces the main data and measuring diagrams of the relatively

large. The out-of-roundness diagrams apply to the plane
noted as C because, according to our experience, the highest out-of-roundness always appears here (Kundrak and Bana,
2003), due to the deforming effect of the three-jaw chuck.
The characteristic out-of-roundness is a three-lobe form for
both processes. Cylindricity error, like out-of-roundness, is
higher for hard turning. The reason for this is the higher
clamping force and form generation with a single-point tool,
whereas in grinding the bore is conical. At the point where
the grinding tool is first applied the diameter is larger, since
on this side the wheel is more overtravelled than on the other
side.
The parallelism is better in hard turning but on the entering
side the diameter is smaller due to insufficient conduction of
the intensive heat. The parallelism error of grinding is higher;
the expansion on the entering side can be explained by the
running out of the end stroke of the wheel to a larger extent.
The final results of measurements, with the addition of
flatness and the axial run-out of face, are presented in Fig. 8.
Although the out-of-roundness is higher in hard turning, it
still satisfies the quality prescribed for IT5. The parallelism
of the generatrices is also suitable. The cylindricity tolerance
is not prescribed, therefore, its generation is less important.
The flatness is adequate for hard turning, but for grinding it is
higher than the permitted limit. The axial run-out is adequate
in both hard turning and grinding but much higher in grinding,
where it is about five times of the hard turned values. These
results show that the ground gear does not fulfil the require-

ments prescribed in the drawing because of its large flatness
error.

The following hard turned and ground pair (Fig. 9) also
possesses a bore with accuracy IT5, but their wall thickness
(rigidity), at the droot /d1 ratio was 1.66, is significantly lower
than the previous one of 2.32. The data for the gears, prescriptions for accuracy, and the diagrams of accuracy is given
in Fig. 9. In comparison with the previous gear, the out-ofroundness is much higher and because of the three-jaw chuck,
a three-lobe form can be observed, which appears here more
clearly than before. The cylindricity error is slightly higher in
hard turning but in grinding the bore diameter is higher on the
entering side. The reason for the conicity is that wheel overtravel occurs to a higher degree here than on the chuck-hand
side. Here the parallelism of the generatrices for hard turning
is also better than in grinding.
In hard turning, at entering a lower accumulation of heat
can be observed, which increases at outgoing. In addition, on
the hard turned piece a softer mark was found on the scanned
side that was abreacted by the single point tool by more material removal. On the bore of the ground gear the result of
unequal overtravel of the wheel also appears.
The results of all performed measurements are summarized in Fig. 9. The out-of-roundness measured in planes A, B,
and C is notable. It is always the highest in plane C because
here the deforming effect of the clamping force is the highest.
Moreover, for this gear neither hard turning nor grinding meet
out-of-roundness limits, especially in the case of hard turning. The prescription for cylindricity is not given, therefore its
shaping is neutral. The parallelism of the bore’s generatrices
is definitely acceptable.
Comparing Figs. 8 and 10, it is prominent that the flatness
error and the axial run-out of the ground part face are 4–5
times higher than those of the hard turned components. This
arises because of the grinding process applied. The face has to
be machined on the internal grinding machine. However, face
grinding can be regarded as a necessary procedure on these
machine tools. The armed apparatus that clamps the grinding

spindle can be brought into work position outside. As a consequence of this we must work with a light apparatus with
low rigidity, which leads to a higher degree of deformation
from the grinding force. Moreover, the working of the appa-

Fig. 10 – Summary of measured results.


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ratus that performs the dressing of the grinding wheel is not
always precise, because sometimes the dressing of the small
cup wheel is done by hand. Because of the effort made for relatively small sizes, the grinding quill is also weak. Therefore
it is no wonder that the face plane is not precise.

337

As for the axial run-out, the surface grinding of the face
is performed on a separate machine in a different clamping.
While the clamping was done on the pitch circle in internal
grinding, for face grinding the clamping is on the opposite
plane surface. The alteration of the clamping surface causes

Fig. 11 – Diagrams of hard-turned and ground profiles.


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cesses, but it is higher in grinding than in hard turning. No

prescription for cylindricity exists for this gear.

7.

Fig. 12 – Summary of measured results.

the significant flatness error and axial run-out. In hard turning such problems do not arise because the whole machining
is done in one clamping, and therefore, the flatness error and
run-out are significantly lower. The out-of-roundness of gear
E belonging to the bore quality IT5, whose magnitude exceeds
the prescribed values, can be explained by the inadequacy of
the clamping device. For such a small wall thickness clamping
with concentrated force is not acceptable. A clamping device
of some other construction is necessary; for this purpose magnetic and hydraulic clamping chucks are the most suitable
options.
Finally we investigated one more hard turned and ground
pair, which agrees with the previous in everything but the bore
diameter’s tolerance, which is not IT5 but IT6. There are no
prescriptions for the flatness and the axial run-out of the face.
The main data of the gear and the diagrams gained during the
measurement can be seen in Fig. 11.
The summary review of the measuring errors can be seen
in Fig. 12. In the bore with quality IT6 the prescribed out-ofroundness cannot be met with hard turning but can be with
grinding. This is because the higher concentrated force necessary for hard turning deforms the workpiece with small wall
thickness to such a degree that after release it possesses outof-roundness higher than that permitted. In grinding, because
of the lower clamping force, lower form error will occur. The
parallelism of the generatrices can be ensured for both pro-

Conclusions

The most critical element in the accuracy of hard turning is
the generation of out-of-roundness. If the workpiece is rigid
enough – droot /d1 ≥ 1.90 – then IT5-level out-of-roundness can
still be achieved but in lower rigidities – droot /d1 < 1.70 – even
values permissible by IT6 cannot be ensured. This is due to the
highly concentrated clamping force from the clamping mechanism with of 12 bar pressure (uncontrollable). This problem
could probably be avoided if, instead of concentrated clamping
force, a magnetic or vacuum chuck working with distributed
force were applied. In this case, the typical three-lobe form
of the bores would disappear. The other prescriptions for the
form and positional accuracy can be ensured by hard turning at IT5 level. In grinding, the flatness and axial run-out of
the faces are critical, and this is about 4–5 times higher than
in hard turning. With the given equipment and technology it
can be hardly reduced. However, conicity on the entering side
can be avoided with the overtravel of the wheel occurring in
lower degree.

Acknowledgement
The authors greatly appreciate the support of the Hungarian
Scientific Research Fund (No. T042962).

references

Barry, J., Byrne, G., 2002. Chip Formation, acoustic emission and
surface white layers in hard machining. Ann. CIRP 51 (1),
65–70.
Kundrak, J., Bana, V., 2003. Geometrical accuracy of machining of
hardened bore holes. In: Fourth Workshop on European
Scientific and Industrial Collaboration, May 28–30, University
of Miskolc, Hungary, pp. 473–480.

Schmidt, J., Dyck, M., 2000. Einstieg in die Trockenbearheitung,
Werkstatt und Betrieb. JAHRG 133, 68–70.
¨
Tonshoff,
H.K., Karpuschewski, B., Borbe, C., 1997. Comparison of
basic mechanism in cutting and grinding hardened steel.
Prod. Eng. 4 (2), 5–8.


Surface & Coatings Technology 203 (2008) 291–299

Contents lists available at ScienceDirect

Surface & Coatings Technology
j o u r n a l h o m e p a g e : w w w. e l s ev i e r. c o m / l o c a t e / s u r f c o a t

The impact of surface integrity by hard turning vs. grinding on fatigue damage
mechanisms in rolling contact
Y.B. Guo ⁎, A.W. Warren
Department of Mechanical Engineering, the University of Alabama, Tuscaloosa, AL 35487, United States

a r t i c l e

i n f o

Article history:
Received 17 March 2008
Accepted 1 September 2008
Available online 11 September 2008
Keywords:

Surface integrity
Hard turning
Grinding
Rolling contact
Fatigue damage

a b s t r a c t
The fundamental knowledge of fatigue damage mechanism is necessary for understanding manufacturing
process effects. The traditional method of artificially created surface defects may accelerate crack propagation and
fatigue. However, the artificial defects will change the surface integrity and therefore alter the nature of fatigue
damage, and thus, the fatigue damage in the presence of artificial defects in literature may not reflect the true
mechanism of real-life fatigue processes. This paper studies the fatigue damage resulting from real-life rolling
contact tests and finite element analysis of AISI 52100 steel and identifies the possible mechanisms for fatigue
failure in the presence of process induced surface integrity. Rolling contact fatigue tests were then performed until
surface spalling had occurred. Surface and subsurface fatigue damage of the samples was then characterized using
optical and scanning electron microscopy (SEM) and surface profiling. Two types of subsurface cracks were
observed: main cracks that propagate parallel to the surface due to subsurface shear stress induced fracture/
debonding of inclusions or second phase particles. Shear stress induced surface cracks propagate at shallow angles
(≈35°) from the surface. Branching cracks eventually form and connect the main crack to surface. The crack SEM
images show that the formation sequence of fatigue cracks is different for the turned and ground surfaces.
Material loss (spalling) occurs as a combined effect of the main, surface, and branching cracks creating small
discontinuous sections of material at the surface which are eventually lost after continued rolling.
© 2008 Elsevier B.V. All rights reserved.

1. Background on rolling contact fatigue damage
1.1. Introduction
Since most mechanical components produced by hard turning and
grinding are widely used in rolling contact applications, the fundamental
knowledge of fatigue damage mechanism is necessary for understanding
manufacturing process effects. Most research to investigate rolling contact

fatigue (RCF) damage were performed by creating artificial surface defects
such as cracks or depressions using Vickers or Knoop indentations. The
purpose of doing this is to accelerate crack propagation and fatigue,
however, creating such artificial defects is not realistic and the results do
not necessarily compare with real-life scenarios. The artificial defects will
change the surface integrity and therefore alter the nature or mechanism
of fatigue damage. Thus, the fatigue damage in the presence of artificial
defects in literature may not reflect the true nature of real-life fatigue
process. Therefore, fatigue tests performed without the presence of
artificial defects (excluding machining profiles resulting from turning and
grinding or minor scratches due to polishing) are necessary and more
representative of real-life applications.
The majority of fatigue mechanisms investigated in other research
is based on the accelerated tests of various materials such as metals and
⁎ Corresponding author. Tel.: +1 205 348 2615; fax: +1 205 348 6419.
E-mail address: (Y.B. Guo).
0257-8972/$ – see front matter © 2008 Elsevier B.V. All rights reserved.
doi:10.1016/j.surfcoat.2008.09.005

ceramics [1–3]. Currently, very little research has been performed to
study fatigue mechanisms of hard machined components. Due to the
very different surface integrity induced by hard turning or grinding,
the fatigue damage mechanisms could be significantly different from
other mechanical components because the fatigue life is very different
for same surface finish [4]. Though a few studies have examined the
general surface integrity factors such as roughness, residual stress, etc
on RCF, the basic limitations is that the conclusions are too general to
explain the specific difference of fatigue performance caused by hard
turning and grinding [4,5].
The objective of the present RCF tests is to study the fatigue damage

mechanisms for crack propagation in real-life tests of samples prepared
by hard turning and grinding. Four types of test samples were prepared:
as turned (AT), turned and polished (TP), as ground (AG), and ground and
polished (GP). Each RCF test was ran until significant fatigue damage (i.e.
surface spalling) had occurred at which time it was stopped. The test
samples were then analyzed in terms of surface and subsurface fatigue
damage considering the effect of process induced surface integrity. A
fatigue damage mechanism has been proposed based on the experimental and simulation results.
1.2. Fatigue crack appearance
Three possible mechanisms for surface fatigue crack growth have
been proposed by Bower [6]: shear (Mode II or III), fluid pressurization


292

Y.B. Guo, A.W. Warren / Surface & Coatings Technology 203 (2008) 291–299

Fig. 1. RCF test setup.

based “wedge effect” [3] (Mode I). The shear model assumes that
cracks are initiated by the Hertzian contact stress field occurring
around the circumference of the contact area. For this model, the fluid
does not have an effect on crack growth. The wedge effect model is
based on crack growth by fluid pressurization generated by the
entrainment of fluid (i.e. lubricant) into the high pressure contact
flowed into the pre-existing crack or surface defect and caused it to
open (Mode I) or fluid entrapment in which closure of the mouth
sealed fluid into the crack and contact motion generates fluid pressure at the crack tip (Mode II). Mode II propagation can be severely
hindered in practice due to friction or interlocking of the crack faces
and the presence of a fluid assists crack growth by reducing friction,

separating the cracks during compressive loading, or by creating
internal pressure. This has been shown by Way [7] who found that RCF
pits only occurred when a lubricant was present.
In rolling contact fatigue (RCF) [1–22], the initial cracks are nearly
always inclined to the surface at a shallow angle (≈20°) to the surface.
For surface originated cracks, most research shows that the orientation
of the crack is highly dependent on the direction of the tractive force. It
is usually observed that the crack grows in a direction opposite the
direction of the tractive force. For subsurface originating cracks, it has
been suggested that the direction of crack growth is dependent on the
direction of rolling [8]. Explanations for the direction of crack propagation have been proposed that consider residual stresses [8,9]
resulting from the machining processes. Residual stresses could open
the crack in the observed orientation while others [9,10] speculate that
the cracks evolve by ductile fracture and that accumulating plastic
deformations control the crack directions. The prediction of a preferred
crack orientation is made more difficult because the topography of the
contacting surfaces is constantly evolving. The changes in surface
topography lead to dramatic differences in the contact pressure and
therefore the contact stresses driving crack propagation. Even small
changes in surface topography lead to a substantial difference in pressure and the presence of surface spalls or pitting can cause the pressure
to greatly exceed the Hertzian pressure [8].

the valleys of one contact surface which is then pressurized by the
other contact surface leading to hydrostatic pressure and crack growth.
The effect of surface roughness on fatigue damage is not fully understood, however, previous research [11] has shown that increasing the
surface roughness above a critical level leads to greater influence of
micro-geometrical features on fatigue strength.
1.3.2. Microstructure effect
The capability of a material to resist fatigue damage during rolling
contact is directly related to its microstructure. Brittle materials such as

ceramics and cast irons usually fail from normal loads, while ductile
materials such as steels fail due to shear stresses. The reason for
materials being either ductile or brittle is directly controlled by the
microstructure and atomic arrangement of the crystal lattice. Previous
research [12] has shown that the presence of a “white layer” induced
by phase transformation on hardened steels produced by abusive
hard turning is detrimental to RCF life for high load applications. A
white layer is formed when cutting temperatures above the austenizing temperatures are achieved and the martensite is then rapidly
quenched. The resulting temperature effects leave an untempered
martensite (UTM) region that is referred to as the white layer. During
rolling contact, the maximum normal stresses are located at the
surface, while maximum shear stresses occur at a certain depth in the
subsurface since friction coefficient is much less than 0.3 for the
well lubricated contacting surfaces. The machined surface of the test
samples could be considered as brittle since the hardness of the heat
treated AISI 52100 steel is 62 HRC. In addition, strain hardening which
occurs during hard turning and grinding increase the hardness in the
near surface by as much as 20%. This is important because it means
the location of crack initiation is dependent on the loading and microstructure of the component. For steels, the origin of cracks is usually at
surface asperities, subsurface inclusions, or cementite–ferrite boundaries [13]. It is important to note that the initial damage does not
always occur at the location of maximum shear stress, but can occur at
any location where the shear stress is greater than the maximum shear
strength of the material.

1.3. Fatigue mechanisms on crack initiation and propagation
1.3.1. Surface roughness effect
Surface roughness is critical to fatigue endurance of components
subjected to RCF. The contact stress between two bodies in contact is
significantly affected by the surface topography of the contact zones.
High surface roughness can cause local contact stresses to be dramatically higher than predictions based on the Hertz theory or finite

element analysis and results in more rapid failure of components.
Surface roughness profiles can also simulate pre-existing surface
flaws such as cracks. The peak and valley profile of turned and ground
surfaces exhibit such a phenomenon and act as notches that experience locally high stress concentrations and are the preferred
locations for crack initiation. During RCF, lubricant can be trapped in

1.3.3. Residual stress effect
Residual stresses can have a significant effect on the development
of fatigue damage [14–20]. Residual stresses are caused by lattice
distortions resulting from the machining process. These lattice distortions serve to increase the strength of the material by hindering
dislocation motion during loading. Compressive residual stress will

Table 1
RCF fatigue test conditions
Load (N) Peak Hertzian stress (MPa) Radius of contact circle (µm) Frequency (Hz)
306

4629

170

173


Y.B. Guo, A.W. Warren / Surface & Coatings Technology 203 (2008) 291–299

293

Fig. 2. Wear track and pitting on the ground and polished sample.

impede crack growth by closing the crack tip, while tensile residual
stress aids crack growth by opening the crack mouth. A mechanical
model [23] has been developed to quantify the effects of the residual
stress and hardness gradient by laser transformation hardening on
crack driving force. During rolling contact applications, the magnitude and distribution of residual stresses is constantly evolving.
Voskamp [17] has shown that residual stresses develop in high carbon
steels (1% C) at sufficiently high loads and these stresses may influence
the direction of crack propagation. Guo et al. [18,19] has shown that
the distinct residual stress profiles by hard turning and grinding only
affect near-surface fatigue damage rather than locations deeper in the
subsurface. The residual stresses affect neither the magnitudes nor the
locations of peak stresses and strains below the surface at high load
applications. The slope and depth of a compressive residual stress
profile are key factors for rolling contact fatigue damage. Equivalent
plastic strain could be a parameter to characterize the relative fatigue
damage.
2. Experimental procedures
Four types of test surfaces of as turned (AT), turned and polished
(TP), as ground (AG), and ground and polished (GP) were prepared.
The AT and AG surfaces have equivalent surface roughness Ra 0.16 µm
and the TP and GP surfaces have Ra 0.06 µm [24]. The experimental
RCF setup shown in Fig. 1 was used to monitor RCF tests of ground
surfaces. This test rig is capable of rotating up to 4000 rpm, thus
allowing the test to run in a reasonable time frame. The load is applied
through the rotating shaft which also houses the slave washer. The
test utilized eight chrome steel balls with the diameter of 5.56 mm
and hardness of 63 HRc. A retainer was used to hold the balls and was
constructed of nylon to minimize AE noises during the test. Slave

washers of the same material as the test samples were created by

turning. An 8 mm radius groove was machined on each slave washer
to maintain the ball's position during operation and provide a lower
contact pressure between the ball and slave washer. The lower contact
pressure in the slave washer would help ensure that fatigue failure
would first occur in the test sample, outside of any other variables. The
testing conditions are shown in Table 1.
In-situ monitoring of fatigue damage was accomplished using a
complete AE acquisition package that is versatile in the application of
fatigue damage analysis. An AE sensor was attached to the test sample
with a holding fixture, while vacuum grease served as the coupling
media between the sensor and test sample. The AE sensor has a
125 kHz resonant frequency and connects to an 18 bit PCI-2 data
acquisition board that was incorporated into a PC. Before the data
reached the PC it was passed through a preamplifier that was set at a
40 dB gain and a threshold of 45 dB was used. The RMS/ASL time
constant for all of the tests was 500 ms with a sampling rate of 10 Hz.
The sampled AE parameters in this study include: absolute energy,
RMS, amplitude, counts, and average frequency.
The test sample is secured in a resting plate which is positioned
with four rods that maintain the rigs position and rigidity. The resting
plate is free to move up and down as it rests on a load cell, allowing
accurate measure of the applied load. An acoustic emission sensor is
mounted directly to the test sample. Prior to startup, a high-temperature
multi-purpose lithium complex grease was evenly distributed across
the test surface, slave washer track, and balls. Additional lubrication
was added at regular intervals throughout the duration of the test.
Rotational speed of the spindle and washer was monitored using an
optical tachometer.
Parallelism and centricity between the slave washer and test
surface is critical to ensure that the pressure distribution is uniform

Fig. 3. Top surface spall on the ground and polished sample.


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Y.B. Guo, A.W. Warren / Surface & Coatings Technology 203 (2008) 291–299

Fig. 4. a. Wear track profile of the turned and polished sample. b. Wear track profile of the as turned sample. c. Wear track profile of the ground and polished sample. d. Wear track
profile of the as ground sample.

across the test surface. For this reason, a dial indicator gage was used
to measure the vertical and horizontal runout of the washer and test
surface. Using this method it was possible to adjust the horizontal and
vertical displacements so that parallelism and centricity would be
achieved.
3. Experimental results and observations
The fatigued samples were analyzed in terms of surface and
subsurface damage. Surface damage reveals itself as pitting and/or

spalling and is formed from an accumulation of cracks that eventually
lead to fracture and material loss. This type of damage can be observed
on the test surface by optical microscopy immediately after the RCF
tests are completed. Subsurface damage is only visible on the crosssection of the test samples and requires cutting and polishing in order
to view the crack characteristics such as depth and orientation.
The significant challenge for this work is to determine the mechanisms for crack growth during rolling contact by analyzing the post
fatigue damage. It is impractical to conduct an experiment of this
nature and document the progression of fatigue damage for every test

Y.B. Guo, A.W. Warren / Surface & Coatings Technology 203 (2008) 291–299

295

Fig. 5. a. Surface spall profile of the turned and polished sample. b. Surface spall profile of the as ground sample.

sample. In order to view the subsurface cracks, the test specimen must
be destructively cross-section which renders the sample useless for
further loading by rolling contact. The technique used in this study
involved subjecting the test specimens to rolling contact until spalling
occurred on the surface. At this time the specimens were removed and
prepared for investigation by cross-sectioning. The fatigue damage
(surface and subsurface cracks) were then observed and characterized
in terms of location, length, and orientation. Comments and speculations were then made regarding fatigue progression based on the
observed fatigue damage. Conclusive or definite statements regarding
the nature of fatigue propagation could not be made due to the lack of
intermediate stages of fatigue damage.
3.1. Surface spalling
Fig. 2 shows a section of the wear track created on the surface of a
GP sample. For brevity, surface damage of each sample is not shown,
but the damage shown is representative of all test samples. Surface
pits and wear debris are scattered throughout the wear track and
become more severe near the spalled region. The width of the wear
track varies from 500 to 750 µm, the larger widths occurring in areas
with the most severe pitting. The wear debris is created from material

Fig. 6. Schematic of test sample preparation.

loss of the ball and test sample asperities (mostly during the run-in
period), and may be responsible for a large portion of the surface pits

by creating dents on the surface as the balls travel across them. Fig. 3
shows an optical image of spalling that was formed on the surface of a
GP test sample. Clearly seen is the substantial loss of material due to
the accumulation of damage that resulted in the observed spalling. On
average, the widths of the surface spalls are between 700 and 850 µm
depending on how long the fatigue damage was allowed to accumulate. Typically, tests are allowed to run for some time after fatigue
damage has been indicated to assure that there is sufficient damage to
investigate. Because the tests run for so long (2–5 weeks depending on
the specific test samples), the post fatigue duration times are not
identical, but are within 1–2 h. For this reason, the width and depth of
the surface spalls can vary to some extent between identical testing
conditions. Radial cracks are clearly seen near the center of the spalls
and are shaped in the form of semicircles that are convex in the
direction of rolling and have an approximate radius of 135–180 µm.
This value is similar to the value of the contact circle radius calculated

Fig. 7. Subsurface fatigue damage of the ground and polished sample.


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Y.B. Guo, A.W. Warren / Surface & Coatings Technology 203 (2008) 291–299

Fig. 8. Subsurface fatigue damage of the as ground sample.
Fig. 11. Finite element simulated of shear stress S23 contour.

load and the interaction of surface topographies of the contacting
surfaces result in cyclic loading frequencies that are much higher than
that of the test surface. For this reason, the balls are predisposed to fail
more rapidly than the other components of the test assembly.

Line traces were made using a Taylor Hobson Talysurf 2000 surface
profiling machine to determine the width, depth, and profile of the
wear tracks and spalled regions on the test samples. Fig. 4 shows
several examples of the wear track profiles. Typically, the wear tracks
have a depth of 2 to 4 µm and range in diameter from 500 to 750 µm.
There is some pile-up of material along the edges of the track that run
parallel to the rolling direction due to plastic deformation (similar to
an indentation) accumulated during the test. Surface spalls are created
when significant material loss has occurred. Fig. 5 shows examples of
spall profiles. These images show the much larger depth (≈60 µm) of
the spalled region when compared to the wear track.
Fig. 9. Subsurface fatigue damage of the turned and polished sample.

3.2. Subsurface cracks
by Hertz theory as shown in [24]. The depth of the spalls ranges from
40 to 60 µm, and is also dependent on the post-failure run time.
An optical image of ball fatigue has been shown in a previous study
[24]. Premature fatigue failure of the balls is due to the kinematics and
loading of the RCF test. Although the balls are the same material and
hardness as the test surface and washer, the load experienced by them
is quite different. While the sample surface and washer experience
cyclic loading at a frequency of ≈173 Hz, the rollers are under constant

Fig. 10. Subsurface fatigue damage of the as turned sample.

Subsurface cracks were investigated by cross-sectioning the test
samples tangential to the wear track as shown in Fig. 6 in a location
where significant fatigue had occurred, mounting in epoxy, and
polishing to a mirror finish using alumina polishing compound. The
samples were also slightly etched using a 2% nital solution to view the

cracks and microstructure in the SEM. The subsurface damage shown
is representative of each testing condition and several observations
can be made.
Figs. 7–10 show the representative subsurface fatigue damage of
the GP, AG, TP, and AT samples, respectively. Upon observing the
collective data, it is identified that there are two distinct crack types.

Fig. 12. Maximum shear stress S23 distribution in the subsurface.


Y.B. Guo, A.W. Warren / Surface & Coatings Technology 203 (2008) 291–299

297

Fig. 13. Orientation of cracks represented by crack analysis tables.

The first is a significant crack that runs parallel to the surface at a
depth of 5.3 to 13.3 µm for nearly the entire length of the specimen
(Fig. 7). This crack is referred to in this paper as the “main” crack. The
other types of cracks are inclined cracks that extend from the surface
to a depth as much as 45 µm. These cracks could extend from the
surface due to high stress concentration induced by sharp surface
asperities for ground surfaces and may be independent of the main
crack. Smaller branching cracks are also observed connecting the
surface to the main crack and it appears as though sufficient branching
cracks in any region causes eventual fracture and loss of material. The
occurrence of a main crack and surface cracks together is not observed
in all test specimens, but at least one or both appear in all fatigued
samples. The mechanisms which form the distinct crack types are not
fully understood, and the purpose of this paper is to explain some

possible causes for the presence of these cracks. Due to the location of
the main crack (5.3 to 13.3 µm below the surface), it is possible that it
is formed due to the magnitude of shear stress experienced by the
sample at this depth which could cause fracture or debonding of
subsurface inclusions or second phase particles located at this depth.
The variation in main crack depth may also be influenced by the
presence (or formation) of residual stresses that would influence the
magnitude/distribution of shear stress experienced by the test
samples. Also, surface topography will greatly influence the distribution of local stresses, and may affect the depth at which the shear
stress is large enough to initiate failure.
A finite element simulation using Abaqus [25] was conducted to
estimate the magnitude and distribution of the stresses resulting from
rolling contact. A 3-dimensional axisymmetric mesh was generated to
simulate the experimental conditions. A rigid roller (d=5.56 mm) was
used to apply the vertical load of 305 N and was given linear and angular
velocities of 3.14 m/s and 1129 rad/s, respectively, which correspond to the
experimental conditions. Shear strength of AISI 52100 steel is approximately 924 MPa [5] and results of the finite element analysis in Fig. 11
show that the depth at which this value occurs is ≈15 µm as shown in
Fig.12. The reason that the observed depth is less than 15 µm is most likely
due to loss of material on the top surface that occurs during the RCF test or
because of stress variations caused by surface asperities. It is important to

note that the FE results assume an ideal flat surface (which is impractical),
and future work will incorporate surface roughness. As the main crack
increases in length, the region above the main crack may behave as a
delaminated surface layer. This will significantly affect the stress
distribution below the main crack because the structural material is no
longer continuous. The delaminated region then becomes subjected to
much higher stress and this may be the cause for formation of the
branching cracks. When a sufficient number of branching cracks are

formed connecting the main crack with the surface, small portions of
surface material eventually are removed by the repeated loading. The
mechanism that causes the formation of surface cracks is quite different.
These cracks appear to form at the surface and travel downward at a
shallow angle to the surface that ranges from 24° to 36° and in a direction
that is usually the same as the direction of rolling. The mechanism for
growth is most likely due to the wedge effect. As the ball approaches, it
forces lubricant into the crack which then propagates due to high pressure
generated as the ball travels across the crack face.
From observation of the fatigue damage, it is apparent that the
formation of the main crack and surface cracks are parallel processes.
However as seen in Figs. 7 and 8, the surface crack must be present
first in order to propagate deeper than the depth of the main crack.
This is because the main crack causes the material to be discontinuous
and surface cracks that form after the main crack forms will arrest
after intersecting it. However, new cracks below the main crack may

Table 3
Crack orientation analysis for case B
Sample
type

θ
(deg)

d1
(µm)

d2
(µm)

Average θ
(deg)

Average d1
(µm)

Average d2
(µm)

Turned and
polished

15
36.5
37
35
30.5
34
33
30
41
39
30
34
47
49
38
9
39

40

15
10.5
7
5.5
12
13
16
19
15
15
14.5
13.5
10.5
10
8.42
17.44
18.46
17.44

7.5
7
2.5
1
4
4
9
7.5
8

14
4.5
5.5
5.5
5
4.21
3.59
15.38
12.31

35.07

12.61

6.07

31.50

15.44

8.87

Table 2
Crack orientation analysis for case A
Sample type

θ (deg)

d1 (µm)

Average θ (deg)

Average d1 (µm)

Turned and polished

43
42.5
44
18.5
16
45
48
26
28
43
40
24
88
41

28
23
24
8
5
15.5
40
5.13
7.18

12.1
16.41
13.33
16.41
9.23

36.71

20.50

27.00

6.16

47.20

13.50

As turned
Ground and polished

Ground and
polished

Table 4
Crack orientation analysis for case C
Sample type

θ (deg)

d1 (µm)

Average θ (deg)

Average d1 (µm)

Turned and polished

23.5
18
39
26

5
23
12.31
10.25

20.75

14.00

32.50

11.28

Ground and polished


298

Y.B. Guo, A.W. Warren / Surface & Coatings Technology 203 (2008) 291–299

Table 5
Crack orientation analysis for case D
Sample
type

θ
(deg)

d1
(µm)

d2
(µm)

Average θ
(deg)

Average d1
(µm)

Average d2
(µm)

As turned

31
15

18
35
40
37
18
25
34
23

46
60
28
32
48
15.38
22.56
26
36.92
56

30
8
14
14.4
16
4.1
11.28
8
24.62
22

21.33

44.67

17.33

30.29

33.84

14.34

Ground and
polished

ble orientations were observed as shown in Fig. 13, but not all occurred
for each test sample. Tables 2–5 represent the data collected for each
crack observed using the SEM. The AG sample did not reveal any
surface cracks and only the main crack was observed so it is not shown
in the tables. The data shows that there is a large variation of crack
depth for each case, which reflects the statistical nature of fatigue.
The variation of local contact stresses due to surface roughness has a
significant effect on the crack propagation mechanisms. It is important
to note however that the average inclination of the cracks to the
surface was relatively uniform (≈35°) regardless of direction of crack
propagation.
4. Fatigue damage mechanism and discussion

be initiated by repeated point loading caused by increased flexibility

of the delaminated near-surface layer. The mechanism for the main
crack is believed to be shear stress. However, the mechanism for
surface cracks is a combination of shear stress (from the compressive
loading of the balls) and tensile stress induced by the presence of
pressurized lubricant that is trapped within the crack as the ball
contact patch travels across the crack mouth. As shown in the finite
element simulation in Fig. 11, the shear stress contour on the top
surface is similar to that of the surface ring cracks observed in Fig. 3.
The formation of surface cracks may be accelerated by tensile stress
concentrations at surface asperities or surface topography profiles
that act as notches. It is believed that the presence of surface cracks
has little effect on the formation of the main crack.
3.3. Depth and orientation of surface initiated cracks
The surface cracks for each test sample were analyzed according
to maximum depth and orientation with the top surface. Four possi-

Fatigue is an accumulation of damage sustained from cyclic stress.
In rolling contact applications, the mechanism of fatigue damage of a
component is dependent on several factors including frequency and
amplitude of loading, microstructure, presence of initial defects,
mechanical properties, and residual stress. For ductile materials such
as steel, fatigue fracture is caused by shear stress. The maximum shear
stress of a component subjected to rolling contact is located in the
subsurface, the depth of which can be estimated by the Hertz theory
or precisely by a finite element analysis.
RCF is a cyclic loading process that is very complex and composed
of many stages. For the current tests, the general progression of fatigue
damage can be outlined as follows:
(1) Initially, the crack free surface is subjected to rolling contact under
specified loading conditions. Immediately after start-up, the

surface experiences elastic and plastic deformation as the balls
travel around the track. The duration of the run-in period is specific
to each test and varies from 4–18 million cycles depending on
several factors including surface roughness, hardness, and load.

Fig. 14. Schematic of RCF fatigue damage mechanism.


Y.B. Guo, A.W. Warren / Surface & Coatings Technology 203 (2008) 291–299

(2) Contact between the surface and balls eventually removes most
of the surface asperities of the test sample and eventually a
wear track is formed that becomes the preferred path of the
balls as they travel across the test surface. The wear particles
that are removed from the surface roughness eventually are
pushed outside of the contact zone, or are displaced onto the
wear track. If a ball travels across a wear particle, it can lead to a
surface depression and eventual ball or surface failure.
(3) The maximum shear stress experienced by the test sample is
located in the subsurface as seen in Figs. 11 and 12, the depth of
which is dependent on the load, contact area, and material
properties. For materials such as AISI 52100 steel with subsurface
inclusions, voids, and interfaces with second phase particles,
fracture and/or delamination can occur at any depth for which
the shear stress is greater than the material shear strength. The
maximum shear stress in the subsurface may be magnified by the
presence of second phase particles and/or inclusions. This may
promote the main crack formation in the subsurface as shown in
Figs. 7–10, especially the as-turned surface in Fig. 10. Because the
depth of the maximum shear stress is constant (for a flat specimen) below the surface of the wear track, a main fracture is often

observed at this depth and can run parallel to the surface
throughout the entire length of the wear track.
(4) The mechanism for inclined or branching cracks nucleation
may be tensile stress located in the vicinity of surface asperities
while propagation occurs due to trapped lubricant that opens
the crack mouth as a result of hydrostatic pressure caused by
the ball as it travels across the crack mouth. The shear stress S23
drives a surface crack propagating downward with an inclined
angle along the rolling direction.
(5) The interaction between an inclined crack and the main crack
usually breaks the inclined crack. The branching cracks from
the main crack may either propagate to or from the top surface.
When a sufficient number of branching cracks reach the
surface, eventual material loss and rapid spalling occurs.
The formation sequence of the main crack and inclined crack may
be different under the influence of distinct surface integrity by turning
vs. grinding. The examination on the SEM images of cracks suggests
two different fatigue damage mechanisms, Fig. 14, for the turned and
ground surfaces. The basic difference is that an initial or main crack
form in the subsurface for turned samples as shown in Fig. 10, while an
initial branching crack could start from ground surfaces and joins with
the subsurface main crack.
5. Conclusions
Fatigue damage of four types of test samples in real-life RCF tests
was analyzed. A fatigue damage mechanism was proposed based on
the experimental and simulation results. The major results can be
summarized as follows:
• Wear tracks created by plastic deformation and material loss from the
rolling process were visible along the circumference of the test sample
surface. Spalling occurs on the sample surface in regions of the wear

track where accumulated fatigue damage results in significant material
loss.

299

• Subsurface damage was characterized by the formation of two distinct
crack formations: the Mode I main cracks that extend parallel to the
surface and surface cracks that are inclined at a shallow angle to the
surface. The locations of main cracks are approximately equivalent to
the depth predicted by a FEA analysis for which the experienced shear
stress was greater than the material shear strength. The main crack is
formed as a result of fracture or delamination of subsurface inclusions
or second phase particle interfaces located at this depth. Possible
explanations for the depth variation in the main crack location were
attributed to differences in residual stress, surface topography, and
statistical variation between the samples created by hard turning or
grinding.
• Surface cracks extend from the surface at shallow angles ≈35°. The
direction of surface crack propagation is somewhat random, but the
majority propagates in the direction of rolling. The mechanism for
surface crack nucleation is tensile stress located at the vicinity of
surface asperities while propagation occurs due to trapped lubricant
that opens the crack mouth as a result of hydrostatic pressure caused
by the ball as it travels across the crack mouth.
• The formation sequence of fatigue cracks is different for the turned
and ground surfaces. The basic difference is that an initial main crack
form in the subsurface for turned samples, while an initial branching
crack could start from ground surfaces and joins with the subsurface
main crack.
Acknowledgement

This research is based upon the work supported by the National
Science Foundation under Grant No. DMI-0447452.
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journal homepage: www.elsevier.com/locate/jmatprotec

A comprehensive characterization of 3D surface topography
induced by hard turning versus grinding
R.A. Waikar, Y.B. Guo ∗
Department of Mechanical Engineering, The University of Alabama, Tuscaloosa, AL 35487, USA

Article history:

Surface topography induced by precision machining is critical for component performance.

Received 3 January 2007

Four representative surface topographies of turned and ground surfaces were prepared

Accepted 30 May 2007

at “extreme” machining conditions (gentle and abusive) and compared in terms of 3dimensional (3D) surface features of amplitude, area and volume, spatial, and hybrid
parameters. The 3D surface topography maps revealed the anisotropic and repeatable nature

Keywords:

of a turned surface which was in sharp contrast with the random and isotropic nature of a

Surface topography

ground surface. In general, a gentle turned surface has higher values of amplitude param-

Machining

eters (arithmetic mean, root mean square, maximum height of summits, maximum depth

Surface integrity

of valleys, and 10-point height) than an abusively turned surface, whereas the opposite was
true for the ground counterparts. Only the gentle ground surface has a negative skewness
which means that the topography distribution is more biased towards the valley side. The
larger kurtosis value of the abusively ground surface implies a more peaked surface topography. The gentle ground and abusively turned surfaces have a much larger bearing area ratio
and therefore better bearing capacity than the gentle turned and abusively ground ones.
The abusively ground surface has higher fluid retainability than other surfaces in terms of
mean void volume. However, surface performance, such as wear and fatigue is dependent
both on surface topography as well as mechanical property and relies on the dominance of
the individual aspect.
© 2007 Elsevier B.V. All rights reserved.

1.

Introduction

Hard turning and grinding are competitive finishing processes
which produce distinct surface topography features due to
the inherent difference in the material removal process. The
single point cutting tool with the defined geometry in turning will naturally produce a more anisotropic surface while
the multiple small abrasives of random geometry in grinding
will produce a more isotropic surface (Malkin, 1989; Marinescu
et al., 2004). The anisotropic turned surface is characterized
by a symmetric and periodic variation of peaks and valleys, whereas the isotropic ground surface shows a more

∗

Corresponding author. Tel.: +1 205 348 2615; fax: +1 205 348 6419.

E-mail address: (Y.B. Guo).
0924-0136/$ – see front matter © 2007 Elsevier B.V. All rights reserved.
doi:10.1016/j.jmatprotec.2007.05.054

unsymmetrical and random distribution of peaks and valleys. The study of surface topography is very essential as it
directly impacts the component functionality, such as friction,
wear, fatigue, and seal behavior. The surface is the boundary between the contact components and its surroundings.
Any interaction between the contact components takes place
at the surfaces. Hence, the characterization of surface topography is a critical factor when component functionality is
concerned.
The geometrical quality is one criterion for the surface
integrity (Guo and Warren, 2004). The state-of-the-art hard
turning can achieve surface values which are, under certain


190

j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 7 ( 2 0 0 8 ) 189–199

conditions, equal to or even better surface roughness than
that by grinding (Pawlus, 1997; Kundrak and Bana, 2003;
Warren and Guo, 2006). The formation of the 3D turned surface topography is the combined action of tool geometry, feed,
depth of cut, material microstructure, and machining system
dynamics. In contrast, the geometries of abrasive particles
and grinding wheel may have more influence on a ground
surface topography. Coarser particles tend to produce rougher
surfaces while fine abrasive particle produces a smoother
surface. A ground surface topography is a function of process
parameters like in turning. However, the relation between
the ground surface topography and process parameters is not

as deterministic as that of a turned one leading to a higher
isotropy.
The bulk surface profile data in literature is limited to 2D
line tracing with a stylus. A 2D profile may not be efficient
in characterizing the surface due to: (a) the reality of contact
surfaces is 3D in nature; (b) the parameter rash or correlation
problem associated with defining 2D parameters; and (c) the
unique effect of 3D surface topography on component functionality. Hence, a 3D surface topography analysis is highly
needed.
Specific to hard turning and grinding, a comprehensive
3D characterization of surface topography generated by hard
turning versus grinding has not yet been done. There is lack of
knowledge of the cause effects of the machining processes on
the 3D surface. The scope of this work is limited to geometrical features of the machined surfaces since the mechanical
and physical properties of turned and ground surfaces have
been well studied (Guo and Sahni, 2004; Guo and Janowski,
2004).
The objective of this paper is to comprehensively characterize and compare 3D surface topography features of the
four types of representative hard turned and ground surfaces
at “extreme” machining conditions (gentle and abusive). This
characterization will shed light on the fundamental relationship between surface topography parameters and functional
aspects.

2.

Background

2.1.
2D surface parameter comparison of turned and
ground surfaces

al., 2002; Renner et al., 2001) with a constant amplitude of the
test piece showed that the specimens which were hard turned
using unworn tools exhibit a higher fatigue strength than the
ground specimens. Only if a massive tool wear (VB = 200 ␮m)
occurs does the fatigue strength of the hard turned test piece
drop below the fatigue strength of the ground ones.
The single cutting edge in turning showed a nearly regular
peak and valley distribution in comparison to grinding with
multiple cutting edges. The peak distance is much wider than
that in grinding. Several studies (Brinksmeier and Giwerzew,
2003; Elbestawi et al., 2003; Seker et al., 2003) have shown that
even if the dimensions of the width of the roughness profile
over the complete measured surface is the same, the distance
and the height between one single peak and valley is completely different in hard turning and grinding.
Both outer and inner hard turning and grinding experiments (Pawlus, 1997; Kundrak and Bana, 2003; Warren and
Guo, 2006; Klocke et al., 2005) have shown that turning at certain machining conditions can achieve equivalent or better
surface roughness Ra and RMS than grinding. The comparison
of other different surface parameters, such as bearing length
curve ratio (Kundrak and Bana, 2003) shows the dissimilarity
between the turned and ground surface features. Even if one
value is the same, both surface topographies cannot be considered as “nearly the same”. Thus, one surface parameter is
not sufficient for characterizing a machined topography.
The skewness values obtained for turned and ground surfaces by Kundrak and Bana (Kundrak and Bana, 2003) were
both negative. Hence, both surfaces showed fluid retention
properties which are good for wear resistance. However, the
hard turned surface had higher negative skewness than the
ground surface which might indicate a better fluid retention
property (Valasek, 1996) and, therefore, better wear resistance
for the hard turned surface. From this point of review, hard

turning would be a recommended finishing process. It should
be realized that the above data was obtained from experiments with a new tool. An increasing tool or grinding wheel
wear will significantly influence the surface topography and a
different set of surface variables are expected. However, very
few data are reported for the cases with worn tools or grinding
wheels.

2.2.

Limitations of 2D surface profile analysis

There are several limitations to a 2D surface profile analysis:
It has shown that hard turning and grinding produce different
surface structures and layers (Guo and Sahni, 2004). Surface structures are consisted of micro-geometrical asperities
and valleys and macro-geometrical features, such as cracks,
undercuts, etc. The surface structures are preferred locations
leading to a stress concentration and therefore crack initiation. Siebel and Gaier (Siebel and Gaier, 1956) have shown that
surface roughness would not influence on the fatigue strength
when the average surface roughness Rz is less than 1 ␮m. However, this characterization of the surface influence by variable
Rz is in adequate, because other surface parameters (Siebel
and Gaier, 1956; Denkena et al., 2002; Borbe, 2001) generated by
hard turning and grinding are very different surface structure.
Even though the notches of the turned and ground surfaces
differ in depth and in occurrence, fatigue tests (Denkena et

(a) As the real nature of contact in surfaces is 3D in nature, a
2D profile can only show the surface roughness at a particular plane. Moreover, the profile may only pass over the
shoulder of a summit and may not represent true peaks
and valleys. If the surface is strictly uniform and the lay
pattern is perpendicular to the plane of the profile, then

the 2D profile may do justice to the surface. However, such
a surface is rare in real world applications. A 2D profile cannot represent an anisotropic nature. Hence, a 2D profile is
inadequate to show the complete and real nature of the
surface. It is from this inadequacy of the 2D parameters
that the need for 3D parameters arises.
(b) The parameter rash (Whitehouse, 1982) problem associated with defining 2D parameters. Parameter rash is


j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 7 ( 2 0 0 8 ) 189–199

191

Table 1 – Hard turning conditions
Turning conditions

Surface type
Turned fresh surface (HTF)

Work material
Cutting tool

Velocity
Feed
Depth of cut (DoC)
Coolant type

AISI 52100 steel
GE BZN 8100 (fresh)
(0.015/15◦ chamfer and
6.35 mm radius)

1.78 m/s
0.0254 mm/rev
0.254 mm
Dry turning

manifested in two aspects. First, parameters defined in
some national and international standards are found to
vary from country to country. Second, the dominance of
manufacturers of measurement instrument on the development of 2D parameters resulted in some problems of
defining parameters. For example, the commonly used Ra
has no direct functional significance and is less significant
in statistics than the root-mean-squares RMS. Third, some
parameters are correlated. A well-known example of this
is the correlation of the arithmetic mean Ra and the Rq .
(c) Most importantly, 3D surface topography has a better correlation with component functionality. The effects of 2D
surface parameters on functionality have been studied
on wear (Suh and Nagao, 1976; Jahanmir and Suh, 1977),
scuffing (Kelley and Lemanski, 1967), pitting (Berthe et
al., 1977), and run in behavior and fatigue (Zhou and
Hashimoto, 1994). To evaluate the influence of surface
topography on component functionality, a set of 3D surface parameters has to be studied to relate them with
functional aspects. General 3D surface parameters have
been defined to explain the fluid retention property of the
surface which could explain potential wear and friction
behavior of the surface (Dong et al., 1994a). The basic difference between the 3D surface parameters between the
turned and ground surfaces has been poorly understood.
Furthermore, if some surface parameters are same for surfaces machined by different processes then explaining the
functional aspects becomes even more difficult.
Although surface topography has been studied in great
detail in terms of 2D parameters, a comprehensive compar-

Turned surface with white layer (HTWL)
AISI 52100 steel
GE BZN 8100 (VB 0.5 mm)
(0.015/15◦ chamfer and 6.35 mm
radius)
2.82 m/s
0.0254 mm/rev
0.254 mm
Dry turning

ison on the nature of 3D surface topography generated by
hard turning and grinding at different machining conditions is
still missing. Additionally, the relationship between 3D surface
topography and functional aspects is poorly understood.

3.

Experiment procedures

Work material of AISI 52100 steel discs of 76.2 mm diameter
and 19.05 mm thickness were heat treated at the austenizing temperature 815 ◦ C 2 h, quenched in an oil bath at 65 ◦ C
for 15 min, and tempered at 176 ◦ C for 2 h producing a final
hardness of 61–62 HRC. The test samples were machined by
both face turning and grinding at machining condition in
Tables 1 and 2.
Since the objective is to study and compare surface topography obtained by turning and grinding a practical approach
is, to generate surfaces at “extreme” machining conditions
including “best” and “worst” achievable machining conditions. Since the machining conditions are very difficult to
define since it depends on the performance of machine tools

and tooling. Consequently, the turning and grinding conditions were divided into two categories, i.e. gentle and abusive
machining conditions in production scenarios. For turning
experiments, the “best” achievable machining condition as
shown in Table 1 is to use a fresh round CBN cutting tool insert
at the gentle machining range to ensure that there would be no
phase transformations on the machined surface. The “worst”
or abusive turning condition was carried out using a worn cutting tool with a flank wear of 0.5 mm and an increased cutting
speed so that a white layer, a surface burn via phase trans-

Table 2 – Grinding conditions
Grinding conditions

Surface type
Ground fresh surface (GF)

Work material
Grinding wheel
Wheel speed
Table speed
Cross feed
Down feed (rough)
Number of passes (rough)
Down feed (finish)
Number of passes (finish)
Coolant type

AISI 52100 steel
Al2 O3 (dia 254 mm)
23.94 m/s
15.24 m/min

1.14 mm/pass
12.7 ␮m/pass
2
5.08 ␮m/pass
1
Water soluble

Ground surface with white layer (GWL)
AISI 52100 steel
Al2 O3 (dia 254 mm)
23.94 m/s
16.62 m/min
1.14 mm/pass
12.7 ␮m/pass
2
26 ␮m/pass
1
No coolant


192

j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 7 ( 2 0 0 8 ) 189–199

Fig. 1 – Fresh surface by hard turning (HTF).
Fig. 3 – Fresh surface by grinding (GF).
formation, appears on the ground surface. The face turning
experiments were conducted using a CNC lathe which maintains a constant cutting velocity to generate uniform surface
integrity.
Gentle grinding operation was performed using an Al2 O3

wheel which was dressed with a diamond wheel prior to the
grinding tests. Ample coolant was used to minimize heat generation or thermal damage on the ground surfaces. Abusive
grinding operation was performed using a dry Al2 O3 wheel
without any coolant with an increased depth of cut and higher
table speed as shown in Table 2.
The 3D surface topography of the machined surfaces was
measured using a Taylor Hobson Talysurf CLI 2000 3D surface
profiling system. Due to the small geometrical features of the
precision machined surfaces, the measurement was carried
out using the inductive gauge with resolution of 10 nm and
measurement range of 2.5 mm. The stylus was used to scan
across a set area of the workpiece. The collected data is processed using the signal processing software package to render
the surface topography maps and 3D surface parameters.

4.

Results and discussion

4.1.

3D surface topography

The 3D topographies of the as machined surfaces are shown
in Figs. 1–4. The turned surface topography in Figs. 1 and 2
shows well-defined peaks and valleys, while the ground ones
in Figs. 3 and 4 show a much more random surface. This
is mainly because of the distinct difference in cutting edge
geometry between hard turning and grinding. Turning uses
a single cutting edge with defined geometry to generate the

Fig. 2 – Hard turned surface with white layer (HTWL).

machined surface and the turning parameters of feed and
cutting edge geometry define the symmetrical topography
of the machined surface. In contrast, grinding is carried out
with a grinding wheel with randomly distributed abrasives of
irregular geometry. The size and shape of the abrasives and
their spacing are the decisive parameters for the resulting
surface topography. Hence, the parameters which affect the
surface topography in turning and grinding are quite different
in nature. The microview of the turned surface is anisotropic
and the ground one is more isotropic.
The functional aspects of a machined surface are not only
dependent on the roughness but also on many other important aspects like the distribution of peaks, the sharpness of
the surface, the bearing area ratio, and other spatial parameters. The 3D topographies of the turned and ground surfaces
clearly show the difference in spacing between the peaks, the
sharpness of the peaks, and the randomness of the profile.
These aspects of the surface can only be best described using
a set of 3D parameters.
It can also be observed that the abusive turned surfaces
show much sharper and random peaks, and surface cracks
than those of the gently turned ones. This aspect of the topography can be measured by the surface bearing area ratio.
Figs. 1 and 2 show the 3D surface topography of the samples produced by gentle and abusive hard turning. The major
differences in process parameter between the two cases are
the higher cutting speed and larger flank wear used for the
abusive turning condition. The higher cutting speed and tool
flank wear cause higher temperatures and more chatter when
compared to the gentle turning condition. The result is the
increased waviness and surface roughness along the cutting
direction for the abusive turned sample. The 2D surface pro-

Fig. 4 – Ground surface with white layer (GWL).


j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 7 ( 2 0 0 8 ) 189–199

193

Fig. 5 – Surface roughness along feed direction of the HTF surface.

Fig. 6 – Surface roughness along cutting direction of the HTF surface.

files of the hard turned surfaces along the cutting direction are
shown in Figs. 6 and 8. It must be noted that all the 2D profiles
have represent pure roughness values, i.e. the waviness components have been filtered out. The profile for the HTF surface
shows a lower roughness value than the profile for the HTWL
surface which is consistent with the 3D topography image.
However, as the feed is constant for both cases the surface
topography is nearly identical with respect to distribution of
peaks and valleys.
Figs. 5 and 6 show the 2D surface profiles of the HTF
and HTWL surfaces along the feed and the cutting directions
respectively. Pt is the maximum peak-to-valley height of the
2D surface profile. The 2D profile along the feed direction
shows a very repeatable form of peaks and valleys whereas the
profile along the cutting direction shows a random nature. The
profile has been taken along the valley portion of the topography. The random nature of the profile may be because of

random processes like machine vibration, surface imperfections of the material, etc. We also notice that the maximum
peak-to-valley height along the feed direction is much greater

than along the cutting direction. This is because the highest
and lowest points of the topography, i.e. the ridges and valleys
formed by turning are captured by trace along the feed direction only. Similar features are demonstrated by Figs. 7 and 8
which represent he HTWL surface.
Figs. 5–8 show that the 2D profile of the HTWL surface has
larger Pt values than those of the HTWL surface along both
feed and cutting directions. This may be because the HTWL
surface has been machined with a worn tool causing deeper
valleys and more vibrations.
Figs. 3 and 4 show the 3D surface topography of the ground
surfaces. The GF surface appears to be more random and
isotropic than the GWL surfaces. Also the surface roughness
of the GF is lower than the GWL surface which may be inferred

Fig. 7 – Surface roughness along feed direction of the HTWL surface.