mechanical considerations - Lehrstuhl für Werkstoffkunde und

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INTERNATIONAL DOCTORAL SEMINAR, May 16 – 19, 2010, Smolenice castle
MICROSTRUCTURE ANALYSIS OF A BANDED
WORK ROLL WITH COMPARISON BETWEEN
DAMAGE APPEARANCE AND THERMOMECHANICAL CONSIDERATIONS
1
1,2
2
3
M. DÜNCKELMEYER , C. KREMPASZKY , E. WERNER , G. HEIN ,
3
K. SCHÖRKHUBER
1
Christian Doppler Laboratory of Material Mechanics of High Performance Alloys,
Technische Universität München; Boltzmannstraße 15, 85747 Garching, Germany
2
Chair of Materials Science and Mechanics of Materials, Technische Universität
München; Boltzmannstraße 15, 85747 Garching, Germany
3
voestalpine Stahl GmbH, voestalpine-Straße 3, 4020 Linz, Austria
ABSTRACT
Banding is a mechanism which degrades the surface of work rolls for hot rolling due to
thermo-mechanical overloads. In order to investigate the cause of banding, material taken
from a banded work roll has been analyzed by light optical microscopy (LOM) and scanning
electron microscopy (SEM) with focus on the transition between banded and un-banded
regions. The mechanisms of crack formation in carbides and in the matrix material are
investigated with respect to the direction of loading. The observed banding depth is compared
with an analytically calculated plastic zone size associated with thermally induced stresses.
Key words: Hot rolling, Banding, Roll damage, Thermo-mechanically induced damage
INTRODUCTION
The performance of work rolls for hot rolling is governed by several degradation
phenomena, such as mechanical fatigue, wear, thermal fatigue, and oxidation. Therefore much
effort has been undertaken to describe and classify thermo-mechanically induced roll damage
like banding [1]. DEBARBADILLO et al. [2] define banding as the surface deterioration of early
finishing stand work rolls brought about by the tearing out of very small pieces of roll
material, see Figure 1. WALMAG et al. [3] report that in a common rolling mill 0.6 % of coils
are rejected due to rolled-in work roll scale which is caused by banding. An optimisation of
the hot rolling process by means of changed process parameters in order to reduce banding is
discussed by ARNCKEN et al. [4]. REIP [5] shows, that the main reason for banding of work rolls
is the high temperature gradient present in the rolling gap. ERICKSON et al. [6] conclude that the
thermo-mechanical fatigue of the roll surface due to continuous temperature cycling and
resulting mechanical stresses are mostly the cause of wear associated with banding. Therefore
the temperature distribution in the work roll und thermally induced stresses during hot rolling
have been investigated by several authors [7-11] with the help of standard finite element (FE)
solvers. DÜNCKELMEYER et al. [12] propose an analytical approach to investigate the instationary
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temperature distribution and the resulting stresses in circumferential direction of the work roll.
Due to its advantages over numerical parameter studies, the approach will be used in this
work in order to correlate experimentally observed damage with theoretically estimated
plastic deformation. Damage is documented with detailed metallographic analyses of the
work roll surface and profile sections of a banded roll segment. Crack formation will be
discussed with respect to damage mechanisms on the microscale.
INVESTIGATED MATERIAL
In order to tolerate high thermal and mechanical loads work rolls of a hot rolling mill
have to fulfill high demands regarding their properties. For this reason mainly High
Chromium Steel (HiCr) as well as High Speed Steel (HSS) rolls are used in the first stands of
a rolling mill. In this contribution a segment of a HiCr roll with a banded surface is
investigated. DEBARBADILLO et al. [2] showed that the sequence of events which characterize
banding failure is the same regardless of the roll material, even though HSS work rolls seem
to withstand banding better than HiCr work rolls.
The chemical composition and the phases present in the microstructure of the
investigated HiCr work roll material are shown in Tables 1 and 2. Primary carbides occur
clustered, whereas secondary carbides are finely distributed in the microstructure.
Table 1 Chemical Composition
C
Si
Mn
Cr
Ni
Mo
V
W
2.82 % 0.63 % 0.95 % 14.84 % 1.32 % 1.13 % 1.15 % 0.04 %
Fig. 1 Surface damage of an
industrially used work roll
Table 2 Phase content (volume fraction)
Martensitic Primary Secondary
matrix
carbides carbides
65 %
22,5 %
12,5 %
Fig. 2 Investigated HiCr roll
segment with banded surface
Fig. 3
Views directions
MICROSTRUCTURAL OBSERVATIONS
Figure 2 shows the cutting pattern applied to the investigated HiCr roll segment. The
specimens are wire-guided electrical discharge machined and then embedded in epoxy-resin.
After grinding and polishing, specimens for light microscopy are etched with Nital.
The transition zone between banded and un-banded areas is investigated in longitudinal,
lateral and top views, see Figure 3. Figure 4 shows an overview of the transition between
banded and un-banded areas in longitudinal view. Carbides (mainly Cr7C3 [13, 14]) appear
dark in the brighter martensitic matrix. Since the work rolls are fabricated by centrifugal
casting, carbides are mainly elongated in radial direction [15]. Near the banded area – which
can be identified macroscopically as non-shining dark surface, see Figure 1 – the un-banded
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area shows surface near damage in form of a crack network. Cracks, perpendicular to the
surface – so called fire cracks - with lengths up to 120 µm can be detected. The thickness of
the removed layer increases until a maximum depth of approximately 150 µm is reached.
Fig. 4 Damage of a work roll in an overview of the transition between banded and un-banded areas in
longitudinal view, SEM
Figure 5 shows carbides fractured parallel to the roll surface in the banded region in
lateral view. A strongly damaged region can be seen in the lower left corner of the image. It
can be imagined that this region is on the verge of its separation from the roll. Figure 6 shows
the transition zone between banded and un-banded regions in top view. Wear marks due to the
contact between the roll and the strip in the rolling gap can be identified in the banded area.
Figure 7 shows the microstructure in longitudinal view. The coalescence of cracks - parallel
and perpendicular to the surface – and the adhesion of the deteriorated layer to the strip in the
rolling gap cause its separation from the roll. A progressed crack network is detected. Figure
8 shows the transition of banded and un-banded region in longitudinal view. All microscopic
observations show that regions with a high density of carbides also exhibit a pronounced
crack network. These regions are predestined for the initial delamination of a thin surface
layer. A fire crack with a depth of 120 µm can be seen in the center of the image. Figure 9
shows the relation between the banded and un-banded roll surface in lateral view. From this
image a maximum banding depth of approximately 150 µm can be measured.
Fig. 5 Carbides fractured parallel to the roll
surface (dark) and martensitic matrix (light), lateral
view, SEM.
Fig. 6 Transition zone between un-banded and
banded areas in top view, SEM.
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carbide
matrix
Fig. 7 Carbides fractured parallel to the roll
surface (light) and martensitic matrix (dark),
longitudinal view. LOM, Nital etching.
Fig. 9:
Fig. 8 Transition zone between un-banded and
banded areas in longitudinal view, SEM.
Banded area in lateral view, LOM
THERMO-MECHANICAL CONSIDERATIONS
The high temperature gradient due to the contact between roll and strip in the rolling
gap causes high residual stresses in surface near layers of the roll and is therefore responsible
for the deterioration of the roll surface [5]. Hence it is useful to know temperatures and
residual stresses in the work roll. In the following an established analytical macroscopic
model [12] is employed in order to correlate an estimated theoretical penetration depth defined as depth to which thermally induced yielding of the work roll material occurs - with
the experimentally observed thickness of removed layers. The analytical model is structured
as follows: Calculation of a) maximum temperature of the roll, b) temperature distribution in
the roll, and c) residual stresses. Stresses due to the rolling pressure in the rolling gap are
neglected since they are 15 times smaller than the thermally induced stresses [5].
The temperature of the roll at the exit of the rolling gap can be described as [16]
Troll
{
(
)}
K1κ 1−0.5T0
h22κ 2t
=
1
−
e
erfc h2 κ 2 t ,
K1κ 1−0.5 + K 2κ 2−0.5
(1)
where indices 1 and 2 denote the strip and the roll, respectively, K and K are the thermal
conductivities, κ and κ are the thermal diffusivities, erfc is the complementary error function,
t is the contact time, T0 is the initial temperature of the strip and the quantity h2 is defined as,
1
1
2
129
2
H ( K1κ 1−0.5 + K 2κ 2−0.5 )
,
h2 =
K1 K 2κ 2−0.5
(2)
where H denotes the thermal resistance of the contact between roll and strip. Since the radius
of curvature of the roll is large compared to the size of the heat affected zone, it is valid to
model the roll in a plane configuration. According to FISCHER et al. [17], the dimensionless
temperature profile can be described using the dimensionless coordinates ξ and ζ . The
circumferential coordinate is given by x=bξ and the radial coordinate is z=ζδ with the heat
penetration depth
bκ 2
(6)
,
δ=
v
where v, b and κ2 denote the rolling speed, the contact length in the rolling gap and the
thermal diffusivity of the roll. The dimensionless temperature profile then is
⎛ ζ2
~
Θ(ξ , ζ ) = ξ exp⎜⎜ −
⎝ 4ξ
⎛ ζ
⎞
π
⎟⎟ −
ζ erfc⎜
⎜2 ξ
⎠ 2
⎝
⎞
⎟.
⎟
⎠
(3)
In the case of heating the roll in the rolling gap the dimensionless temperature of the roll is
~
(4)
Θ(ξ , ζ ) = Θ(ξ , ζ )
for 0 ≤ ξ ≤ 1,
whereas in the cooling zone the corresponding temperature is given by
~
~
(5)
Θ(ξ , ζ ) = Θ(ξ , ζ ) − Θ(ξ − 1, ζ )
for ξ ≥ 1.
A multiplication of the dimensionless temperature function Θ with the maximum temperature,
Troll, yields the temperature distribution. The resulting stresses are calculated via constitutive
laws [12]. In the linear elastic case the tangential and axial stress components can be
estimated by [18]
Eα ∆T
(7)
σ =−
,
1−ν
where E, ν and α denote Young’s modulus, Poisson’s ratio and the coefficient of thermal
expansion, respectively. Plastic deformations are displayed with a kinematic hardening law
according to LEMAITRE and CHABOCHE [19]. The flow criterion is
(8)
σ − q − Y (T ,η) = 0,
where q is the backstress and Y is the yield strength as a function of the temperature T and
loading direction η (tension: 1, compression: -1). In the following, typical rolling mill values
are taken for kinematical, geometrical and thermo-mechanical data [5]. Material data for a
HiCr roll are taken from [5, 12, 14, 15].
Figure 10 shows the calculated temperature profile of the roll. Heating in the rolling
gap causes a strong increase of the temperature within a thin surface layer of the roll. No
noticeable temperature rise can be observed for depths greater than 3 mm. The analytically
obtained temperature profile at the exit of the rolling gap and the narrow zone of thermally
induced yielding can be seen in Figure 11. Thermal strains cause thermally induced yielding
in layers as thick as 180 µm. Deeper into the roll, the strains are purely elastic.
130
Fig. 10 Analytically obtained temperature
profile.
Fig. 11 Analytically obtained temperature profile at
the exit of the rolling gap. Depth range of thermally
induced yielding of the work roll material.
DISCUSSION
In the following the investigated damage of the microstructure is compared with the
analytical predictions. It is established that banding is mainly triggered by cyclic thermal
loading [5] and that it occurs in areas with a high density of fractured carbides. Therefore, the
individual phases of the microstructure and the mechanisms of banding on the micro level are
investigated.
Figure 12 shows the stress-strain curve calculated from eqs.(1)-(8) for one heating /
cooling cycle together with a schematic representation of crack nucleation and propagation. It
can be seen that sufficiently high compressive loads in circumferential direction induce short
cracks parallel to the roll surface within carbides. Tensile loads in circumferential direction
lead to cracks perpendicular to the roll surface within the matrix [5]. Even though the yield
strength for tensile loading is not reached in this example, kinematic hardening (Bauschingereffect) of the material may lead to a decreasing yield strength which may be surpassed during
subsequent cycles. Through several repetitions of this cycle a crack network is formed. It is
assumed that the damaged layer is delaminated by adhesion between the strip and the work
roll or the back-up roll and the work roll. This assumption is supported by the experimentally
observed transition regions between banded and un-banded areas, compare Figure 4. If the
macroscopical mechanical properties of the work roll material are compared with those of its
constituting phases (matrix, carbide inclusions), it can be found that the matrix starts to flow
during the heat input in the rolling gap. Carbides can be regarded as elastic until their brittle
failure [5]. Figure 13 illustrates the shear stress acting on the carbide causing its fracture: Due
to the heat input in the rolling gap the material starts to flow plastically. The movement of the
material is impeded by the geometry of the roll and the small heat affected layer. Only an
expansion in radial direction is possible, whereby the matrix material starts to flow around the
carbide. The stresses transferred into the carbide lead to the fracture.
Figure 5 shows carbides fractured parallel to the roll surface, which confirms the
assumptions made above. All microscopic observations in banded areas show the occurrence
of fractured carbides and fire cracks up to depths of 150 µm. This correlates well with the
analytically calculated thermally induced yielding zone of 180 µm, see Figure 10.
Figure 6 shows surface deterioration with carbides broken out in the banded zone. This
corresponds to decohesion of a pre-damaged layer due to adhesive contact in the rolling gap.
131
Figures 7 and 8 show coalesced crack networks. Here, carbides broken perpendicularly to the
surface can be seen. Figure 8 shows the transition zone between banded and un-banded areas.
A strong deterioration of the transition zone can be observed. Cracks penetrate into the
material up to a depth of 150 µm.
Fig. 12 Stress-strain curve of one heating / cooling cycle.
Depiction of crucial loading states inducing cracks
Fig. 13 Simplified sketch of the shear
stress inducing fracture of a carbide
CONCLUSIONS AND OUTLOOK
An analytical model is set up in order to estimate strains and residual stresses in
circumferential direction near the surface of a work roll and to correlate these with
microscopical observations. Carbides mainly show cracks parallel to the roll surface, caused
by a compressive load due to the high temperature gradient in the rolling gap. Crack
propagation normal to the work roll surface is assumed to be driven by thermo-mechanical
loads induced by high temperature gradients in surface near layers during cooling of the roll.
Distributed short cracks of carbides and their coalescence with perpendicular cracks in the
matrix (fire cracks) combined with decohesion in the rolling gap are the main failure
mechanisms. Fire cracks in areas without broken carbides are not critical regarding banding
damage. It can be observed additionally that the maximum thickness of the worn layers does
not exceed 150 µm, which corresponds with the extent of the analytically calculated plastic
zone. This work complements the efforts of DEBARBADILLO et al. [2] and ARNCKEN et al. [4] and
can be seen as an extension with respect to the damage mechanisms on the microscale.
According to ARNCKEN et al. [4] a reduction of banding during the hot rolling process could be
achieved through water cooling of the strip, shortly before it enters the rolling gap.
DÜNCKELMEYER et al. [12] propose a well-considered selection of the rolling speed in order to
reduce banding. On the basis of this work guidelines can be set up to reduce banding from a
microstructural point of view: Small, equiaxed and statistically uniformly distributed carbides
(as is the case in HSS rolls) could reduce the banding tendency of HiCr work roll. However,
currently this realisation is hardly possible due to the established production process of the
HiCr work rolls.
132
ACKNOWLEDGMENTS
The authors gratefully acknowledge the financial support of the Christian Doppler
Research Society (CDG).
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133
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