Paulo D.J. Surface Integrity in Machining
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3 Residual Stresses and Microstructural Modifications 111
Thus, Lindsay and Hahn [46] found by measurements on the grinding wheel sur-
face that in fine grinding only 1−3% of grinding grains enter into real contact with
the workpiece material. The contact surface between the grinding grains and the
workpiece surface represents into the grinding tool and usually a larger part into the
workpiece material. Inadequate selection of the machining conditions with respect
to the given workpiece material and tool material may lead to too a strong heat flow,
resulting in more or less distinct changes or damage of both materials in the pair.
In the analysis of microstructural changes in a thin surface layer of the material,
it is possible to evaluate the size of the heat-affected zone. In Figure 3.59 we can
see the time variation of the maximum temperature on the surface at the particular
depths with respect to the workpiece speed v
w
.
Knowing the temperature of melting and the temperature of the austenitization
of the discussed steel, we can define the depth of the re-melted layer and the depth
of the heat-affected zone.
Figures 3.60 and 3.61 show the effects of microstructural changes in the heat-
affected zone due to the maximum temperatures on the surface and in the surface
layer. The changes in the microstructure, microhardness and residual stresses ex-
tend to the depth of austenitization.
Figure 3.58. Geometry of theoretical contact and real contact area in grinding. The elastic
flattening of the wheel causes an increase of interaction length [46]
Figure 3.59. Maximum temperature drop as a function of depth in base material at various
work speeds v
w
[42]
112 J. Grum
For the grinding process the following conditions have been selected:
• different kinds of grinding wheels;
• different grinding conditions (fine, medium fine and rough).
Under different machining conditions of grinding, different temperature cycles
were obtained on the surface and in the depth of the heat-affected zone that has
effected microstructural changes and changes in the microhardness and residual
stresses. Thus, on the surface a maximum temperature higher than the melting
temperature of the workpiece material was obtained. The depth of the re-melted
layer is only a few micrometers and makes a very fine ledeburite microstructure
containing fine cementite spread in residual austenite. The newly formed micro-
Figure 3.60. Grinding temperature cycles on the surface with respect of melting and aus-
tenitization temperature of the material [42]
Figure 3.61. Surface damage of a ground surface in terms of microstructure change, micro-
hardness variation and residual stress variation [42, 46]
3 Residual Stresses and Microstructural Modifications 113
structure has a slightly lower hardness than martensite. The residual stresses in the
thin surface layer will be tensile due to plastic deformation of the surface layer in
grinding caused by tensile forces in the contact zone of the workpiece material and
to this should be added also the tensile stresses induced by the occurrence of resid-
ual austenite. In medium–fine grinding conditions, on bearing location A relative
grinding tensile stresses amount to +425
N/mm
2
and then change the sign in the
depth around 175
μm (Figure 3.62(a)). the relative grinding stress is obtained by
measuring the residual stress after induction surface hardening and then by measur-
ing on the same spot after induction surface hardening and grinding and then calcu-
lating their difference In Figure 3.62(b) we can see the measured absolute residual
stress profile after induction hardening and grinding, and the measured residual
stress profile after induction hardening. The results confirm as predominant the
residual stresses, stresses induced by the plastic deformation of the material and
a lesser influence of tensile stresses caused by the formation of residual austenite.
On the basis of the measurements of residual internal stresses after induction hard-
ening and/or induction hardening and grinding, we can conclude that:
For residual stresses after hardening and grinding, the conditions of rough-
grinding are a more favorable choice. They lower to a lesser extent the desirable
compressive residual stresses after induction hardening.
Grinding conditions can be chosen also so that the melting temperature of the
workpiece material (fine-grinding conditions) is not exceeded. Then, the favorable
compressive stresses after induction hardening are lowered only due to plastic
deformation of the workpiece material during the process and thus relatively low
tensile residual stresses are obtained. However, we should take into account that
this will significantly lower the productivity.
Special attention should be paid to the selection of the kind of grinding wheel in
terms of grinding wheel material, binding agent, hardness and pore density, since
by a right selection we can contribute to higher cutting efficiency concerning the
plastic deformation of the workpiece material. In this way we can keep the grind-
ing tensile stresses as low as possible and make the compressive residual stresses
induced by induction hardening the prevailing kind.
(a) (b)
Figure 3.62. Residual-stress profile after induction-hardening and absolute grinding stress
(a) and the profile of relative grinding stress (b) on bearing location A [43]
114 J. Grum
Induction hardening of surfaces creates a very desirable residual stress state. Re-
sidual internal stresses are always of compressive nature and are usually present to
the depth of the induction-hardened layer [45].
A major difficulty in induction hardening is, however, to ensure a very slight/
slow variation in hardness and the existence of compressive residual stresses in
transition areas to the hardness of the core material. By gently varying the hardness
and existence of compressive stresses in the transition area it is possible to dimin-
ish the notch effect induced by stress concentration. Additional grinding of an
induction-hardened surface deteriorates the stress state in the surface layer, since
grinding has always induced tensile stresses. By a correct selection of machining
conditions and grinding wheel, taking into account its properties, the engineer will
contribute to lower tensile residual stresses and will avoid deteriorating the favor-
able internal stress state after induction hardening.
Xu [49] studied the effect of grinding temperature on the surface integrity of
a nickel-based superalloy. Excellent properties of superalloys together with their
poor thermal diffusivity make them extremely difficult to machine and may lead to
elevated temperatures at the grinding zone and possible thermal damage to the
workpiece during grinding with abrasive wheels. K417 is a relatively new kind of
turbine vane materials, which is a cast nickel-based superalloy and features good
comprehensive properties such as high-temperature strength, high corrosion resis-
tance, and good fatigue resistance.
Figure 3.64 shows residual-stress profiles in ground surfaces having different
colors in which only the stresses in the direction perpendicular to the grinding
direction are given. It can be seen that compressive stresses were generated on the
ground workpiece surface layer.
Away from the ground workpiece surface, the compressive stresses increase to a
maximum value and then decreases to that of the base material. The maximum
stress on the puce surface is slightly smaller than for the other two cases and the
location of the maximum stress is deeper from the ground surface.
Figure 3.63. Residual-stress profile after induction hardening and grinding (absolute stress)
and the profile of relative grinding stress on bearing location G [43]
3 Residual Stresses and Microstructural Modifications 115
Figure 3.64. Residual-stress profiles at different grinding conditions showing various colors
on the surface [49]
3.7 Modeling of Thermally Induced Damage in Grinding
Snoeys et al. [50] studied thermal models yielding a reliable evaluation of grinding
processes. The relationship of grinding temperatures in the contact zone depending
on kinetic parameters can be determined. The depth of the heat-affected zone in the
workpiece can be calculated, workpiece burn predicted and the cause of residual
stresses distribution can be calculated.
Their results show that grinders may lead to differences in heat balance and
heat-flux distribution regarding reduced thermal damage. Maris and Snoeys [47]
developed a reliable thermal model for thermal distribution prediction in workpiece
during conventional grinding process at band heat source, depending on workpiece
velocity v
w
.
Figure 3.65 shows a schematic presentation of the various aspects of thermal
damage in grinding as proposed by König [51]. Thermal damage is understood as
a modification of the physical chemical properties of the workpiece surface layer.
It can be manifested as the depth of heat-affected zone with untempered marten-
sitic layer, oxidized layer, burned surface area, residual stress profile, surface
cracks increased sensitivity for stress corrosion, reduced fatigue resistance, etc.
The knowledge of the grinding temperature distribution is therefore essential for
predicting or eliminating thermal damage [48].
A two-dimensional model can be used, provided that the grinding width is large
with respect to the contact length (Figure 3.66).
The model itself is characterized by three physical quantities:
• the heat input into the workpiece and its distribution over the contact area;
• the real contact length (l
c
); and
• the thermal characteristics of the workpiece material.
a) Actual heat input
The fraction of the heat source flowing directly into the workpiece is of special
importance for its thermal load. Roughly 60−95% of the total generated energy
flows into the workpiece, yielding a fast local increase of the temperature field.
116 J. Grum
b) Contact length
Another essential parameter is the real contact length l
c
, which is a function of the
wheel dimensions, compositions and characteristics.
The dimensions of the length of wheel−work contact (l
c
) depends on the speeds
and feed parameter because they govern the normal grinding force component in
the contact area.
c) Thermal characteristics
Yet other physical quantities are the thermal characteristics of the workpiece and
the grain material. Thermal characteristics are temperature dependent, using an
adequately chosen average value may give a fairly good approximation of the tem-
perature distribution. This assumption, of course, allows for an important simplifi-
cation of temperature calculations.
Information related to the thermal and mechanical characteristics may be found
in databanks.
Figure 3.65. Thermal influence upon ground workpiece surface [51]
Figure 3.66. Two-dimensional thermal model representing grinding action [50]
3 Residual Stresses and Microstructural Modifications 117
It shows a theoretical evaluation of the depth of these influenced zones (Fig-
ure 3.67). The three critical temperatures are:
• tempering at 150°C;
• martensitic at 250°C;
• austenitic transformation at 800°C, for this particular steel (100Cr6), define
three depths of affected zones;
• the heat affected zone z
T
to be compared with the hardness affected zone z
H
;
• the stress-affected zone z
MS
associated with the depth at which the residual
stresses are zero z
s
;
• the untempered martensite layer z
UM
corresponding with the trans-formation
zone z
ac1
.
It has been found that the depth of the tensile residual stress zone for hardened
and tempered 100Cr6 ball-bearing steel can be calculated using the martensitic
transformation temperature as a crucial temperature.
Mechanism of inducing residual stresses. The origin of thermally induced ten-
sile stresses during grinding processes is schematically illustrated in Figure 3.68.
The thermal cycle (0−ΔT) imposed upon the uppermost layer would result in
a large thermal expansion (α ΔT); due to the deformation constraints created by the
much stiffer and not thermally affected of the workpiece material large compres-
sive stresses are built up during the heat-up period.
Figure 3.67. Thickness of thermal damaged layer: the three critical temperatures 150°C,
250°C and 800°C define three depths of affected zones [50]
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The compressive stresses may easily surpass the yield stress at elevated tem-
peratures. During the cooling-down period the stresses in the upper layers change
elastically the stress−strain diagram. At the end of this thermal cycle tensile resid-
ual stresses are created that often are close to the yield of the abrasive material.
The stress−strain cycles of the upper layer, taking into account changes of mate-
rial characteristics (yield, Young’s modulus) due to temperature effects and strain
hardening are represented in Figure 3.69.
Figure 3.70 indicates typical residual-stress profiles below the surface. As indi-
cated on this figure the kinematical parameters on residual stresses turn out to be
quite similar as their effect upon the temperature, magnitude in the contact zone.
Contrary to thermally induced stresses, mechanical-induced stresses are mainly
yielding compressive stresses because of the Hertzian type of the stress field in the
vicinity of the contact points with the grain tips.
Figure 3.68. Mechanism of thermally induced tensile stresses in grinding [50]
Figure 3.69. Stress−strain cycle of upper layer taking into account changes of material
characteristics due to temperature effects [48]
3 Residual Stresses and Microstructural Modifications 119
As the mechanical effects of the grinding wheel are limited to depths of the or-
der of 5 to 10
µm and considering the induced stresses, it means that the tempera-
ture history is the main cause of the residual stresses in grinding.
This effect is most important at the surface of abusively ground workpieces of
easy-to-quench steels.
Compressive stresses occur due to the formation of untempered martensite and
also to a second plastification of this layer.
In order to induce compressive stresses by “gentle” grinding operations on con-
dition that the generated amount of heat is low enough and at sufficient workspeed
to ensure that the temperature is always below the transformation. In such cases,
thermal yielding can be avoided and mechanically induced stresses may become
dominating. All the factors yielding low workpiece temperatures reduce tensile
residual stresses.
Figure 3.71 shows the relationship between observed tensile residual stresses
and calculated maximum temperature.
This relationship is given for alloyed high carbon steels and low carbon steels
indicates the dominating role of temperature effects upon the residual stresses.
Figure 3.70. Residual stresses after plunge grinding of 0.3% carbon steel (hardened) [50]
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In conclusion, the thermal damage of the workpiece can be predicted according
to appropriate thermal model at the critical temperature yielding overtempering and
austenitization. Experiments show a good correlation between predicted and meas-
ured depths of heat affected zones.
The residual stresses are a direct consequence of the high temperature gener-
ated. The resulting tensile stresses are due to plastic deformation caused by thermal
expansion when grinding hardened steels with high martensite content.
The fatigue strength of ground workpieces is reduced by surface microcrack
generation and tensile residual stresses. An adequate choice of grinding parameters
during finishing, especially large workspeeds, may reduce this danger significantly.
Moulik et al. [52] presented simulation of thermal stresses on grinding.
An efficient finite-element procedure has been developed to calculate the tem-
peratures and stresses arising due to a moving source of heat. The procedure is
applied to calculate the thermal stresses produced in hardened steels during grind-
ing. The thermal load during grinding is modeled as a uniformly or triangularly
distributed, 2D heat source moving across the surface of a half-space, which is
insulated or subjected to convective cooling. Figures 3.72(a) and (b) show the
calculated temperature distributions across the surface of the specimen assumed to
be an elastic solid when the heat source has moved.
The results of Figure 3.72(a) correspond to a uniformly distributed heat source,
while those of Figure 3.72(b) are for the case of a triangularly distributed heat
source having the same amount of total heat generation. Maximum surface tem-
peratures produced by both types of heat sources are about the same at approxi-
mately 580°C but their locations are different. The peak surface temperature is
shifted much more towards the trailing edge of the source for the triangular heat
source than for the uniformly distributed one.
In order to calculate the residual stresses produced by the moving heat source, it
is necessary for the specimen to be cooled down to room temperature. This occurs
in grinding due to heat losses through the free surfaces of the specimen. The finite-
element procedure was used to calculate the residual-stress profiles in the elastic-
Figure 3.71. Relationship of temperature effects (thermal expansion and microstructural
changes) and residual tensile stresses [50]