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NSC-F813
DESCRIPTION
DUAL-PHASE TYPE HIGH-STRENGTH STEEL SHEETS HAVING
HIGH IMPACT ENERGY ABSORPTION PROPERTIES AND
A METHOD OF PRODUCING THE SAME
Technical FiE~ld
The preaent invention relates to dual-phase type
high-strength steel sheets, for automobiles use, which
have excellent dynamic deformation properties and exhibit
excellent impact absorption properties, and are intended
to be used as structural members and reinforcing
materials primarily for automobiles, as well as to a
method of producing them.
Background Art
The app:Lications of high-strength steels have been
increasing for the purpose of achieving lighter weight
vehicle bodies in consideration of fuel consumption
restrictions on automobiles and even more applications
for high-strength steel are expected as domestic and
foreign restrictions, relating to estimated impact
absorption properties in automobile accidents, become
rapidly more broad and strict. For example, for frontal
collisions of passenger cars, the use of materials with
high impact absorption properties for members known as
"front side :members" can allow impact energy to be
absorbed through collapse of the member, thus lessening
the impact experienced by passengers.
However, conventional high-strength steels have been
developed with a main view toward improving press
formability, and doubts exist as to their application in
terms of impact absorption properties. Prior art
techniques relating to automobile steel with excellent
impact absorption properties and methods of producing it
have been developed which result in increased yield
strength of steel sheets under high deformation speeds as
an indicator of impact absorption properties, as
disclosed in Japanese Unexamined Patent Publication No.
7-18372, but because the members undergo deformation
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during the shaping process or during collision
deformation, it is necessary to include a work-hardening
aspect to them yield strength as an indicator of impact
resistance, and this is inadequate in terms of anti-
s collision sa:ety in the prior art described above.
In addil:ion, since the strain rate undergone by each
location upon automobile collision reaches about 103
(s 1), consid~aration of the impact absorption properties
of the mater_~als requires an understanding of the dynamic
deformation properties in such a high strain rate range.
Also, high-si~rength steel sheets with excellent dynamic
deformation properties are understood to be important for
achieving both lighter weight and improved impact
absorption p:coperties for automobiles, and recent reports
have highlighted this fact. For example, the present
inventors have reported on the high strain rate
properties and impact energy absorption properties of
high-strength thin steel sheets in CAMP-ISIJ Vol.9
(1966), pp.1.112-1115, wherein they explain that the
dynamic strength at a high strain rate of 103 (s 1)
increases dramatically compared to the static strength at
a low strain rate speed of 10-3 (s-1), that absorption
energy durin~~ crashes is increased by greater steel
material strengths, that the strain rate dependency of
materials depends on the structure of the steel, and that
TRIP type steel (Transformation induced plasticity type
steel) and dual-phase (hereunder, "DP") type steel
exhibit both excellent press formability and high impact
absorption properties. Also, the present inventors have
already filed Japanese Patent Applications No.8-98000 and
No.8-109224 relating to such a DP-type steel, among which
there are proposed high-strength steel sheets with higher
dynamic strength than static strength, which are suitable
for achieving both lighter weights and improved impact
absorption properties for automobiles, and a process for
their production.
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As mentioned above, although the dynamic deformation
properties oi: high-strength steel sheets are understood
at the high strain rates of automobile collisions, it is
still unclear what properties should be maximized for
automobile members with impact energy absorption
properties, and on what criteria the selection of
materials should be based. In addition, the automobile
members are produced by press forming of steel sheets,
and collision impacts are applied to these press formed
members. However, high-strength steel sheets with
excellent dynamic deformation properties as actual
members, bas<~d on an understanding of the impact energy
absorption p=roperties after such press forming, are still
unknown.
For press forming of members for collision safety, a
combination of excellent shape fixability, excellent
stretchability (tensile strength x total elongation >_
18,000) and ~axcellent flangeability (hole expansion ratio
<_ 1.2) is desirable, but at the current time no material
has provided both excellent impact absorption properties
and excellent press formability.
Disclosure of the Invention
The present invention has been proposed as a means
of overcoming the problems described above, and provides
dual-phase type high-strength steel sheets for
automobiles use, which have excellent impact absorption
properties and excellent dynamic deformation properties,
as well as a method of producing them.
The invention further provides dual-phase type high-
strength steel sheets, for automobiles, with excellent
dynamic deformation properties, which are high-strength
steel sheets used for automotive parts, such as front
side members, and which are selected based on exact
properties and standards for impact energy absorption
during collisions and can reliably provide guaranteed
safety, as well as a method of producing them.
The invention still further provides dual-phase type
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high-strength steel sheets for automobiles with excellent
dynamic deformation properties, which exhibit all the
properties suitable for press forming of members,
including exc:ellent shape fixability, excellent
stretchabilii~y and excellent flangeability, as well as a
method of producing them.
The invE~ntion was devised to achieve the objects
stated above by the following concrete means.
(1) A dual-phase type high-strength steel sheets
having high impact energy absorption properties,
characterized in that the final microstructure of the
steel sheet _i.s a composite microstructure wherein the
dominating phase is ferrite, and the second phase is
another low temperature product phase containing
martensite a1= a volume fraction between 3% and 50% after
deformation ~3t 5% equivalent strain of the steel sheet,
wherein the difference between the quasi-static
deformation strength 6s when deformed in a strain rate
range of 5 x 104 - 5 x 103 (s 1) after pre-deformation of
more than 0% and less than or equal to 10% of equivalent
strain, and the dynamic deformation strength od when
deformed in ~3 strain rate range of 5 x 10z - 5 x 103 ( s 1 )
after the aforementioned pre-deformation, i.e. (od - 6s),
is at least GO MPa, and the work hardening coefficient at
510% strain is at least 0.13.
(2) A dual-phase type high-strength steel sheet
having high .impact energy absorption properties,
characterized in that the final microstructure of the
steel sheet is a composite microstructure wherein the
dominating phase is ferrite, and the second phase is
another low temperature product phase containing
martensite at a volume fraction between 3% and 50% after
deformation at 5% equivalent strain of the steel sheet,
wherein the average value odyn (MPa) of the deformation
stress in th~~ range of 310% of equivalent strain when
deformed in .a strain rate range of 5 x 102 - 5 x 103
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(s-1), after pre-deformation of more than 0~ and less
than or equa7_ to 10~ of equivalent strain, satisfies the
inequality: cfdyn ? 0.766 x TS + 250 as expressed in terms
of the tensi7_e strength TS (MPa) in the quasi-static
tensile test as measured in a strain rate range of 5 x
104 - 5 x 103 (s-1) prior to pre-deformation, and the
work hardening coefficient at 5~10~ strain is at least
0.13.
(3) A dual-phase type high-strength steel sheet
having high impact energy absorption properties according
to (1) or (2) above, characterized in that the ratio
between the yield strength YS(0) and the tensile strength
TS'(5) in them tensile test after pre-deformation at 50 of
equivalent si~rain or after further bake hardening
treatment (BH treatment) satisfies the inequality
YS(0)/TS'(5) <- 0.7, and also satisfies the inequality:
yield strength YS(0) x work hardening coefficient ? 70.
(4) A dual-phase type high-strength steel sheet
having high :impact energy absorption properties according
to any of (1), (2) or (3) above, characterized in that
the average grain size of the martensite is 5 ~m or less,
and the average grain size of the ferrite is 10 ~.m or
less.
(5) A dual-phase type high-strength steel sheet
having high .impact energy absorption properties according
to any of (1), (2), (3) or (4) above, characterized by
satisfying t:he inequality: tensile strength (MPa) x total
elongation (~) ? 18,000, and by satisfying the
inequality: hole expansion ratio (d/do) >_ 1.2.
(6) A dual-phase type high-strength steel sheet
having high impact energy absorption properties according
to any of (1), (2), (3), (4) or (5) above, characterized
in that the plastic deformation (T) by either or both a
tempering rolling and a tension leveller satisfies the
following inequality.
2.5 fYS(0)/TS'(5) - 0.5} + 15 >_ T ? 2.5
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~YS(0)/~'S'(5) - 0.5} + 0.5
(7) The dual-phase type high-strength steel sheet
having high impact energy absorption~properties according
to the invention is also a dual-phase type high-strength
steel sheet with excellent dynamic deformation properties
according to (1) to (6) above, characterized in that the
chemical compositions, in terms of weight percentage, C
at 0.02~0.25~;, either or both Mn and Cr at a total of
0.15~3.5~, one or more from among Si, A1 and P at a total
of 0.02~4.0~, if necessary one or more from among Ni, Cu
and Mo at a total of no more than 3.5~, one or more from
among Nb, Ti and V at no more than 0.30, and either or
both Ca and REM at 0.0005~0.01~ for Ca and 0.005~0.05~
for REM, with the remainder Fe as the primary component.
(8) The dual-phase type high-strength steel sheet
having high impact energy absorption properties according
to the invention is also a dual-phase type high-strength
steel sheet with excellent dynamic deformation properties
according to (1) to (7) above, characterized in that one
or more from among B (__<0.01), S (__<0.01%) and N (<_0.02~)
are further added if necessary to the steel.
(9) The method of producing a dual-phase type high-
strength hot--rolled steel sheet having high impact energy
absorption properties according to the invention is a
method of producing a dual-phase type high strength hot-
rolled steel sheet with excellent dynamic deformation
properties according to (1) to (8) above, characterized
in that after a continuous casting slab is fed directly
from casting to a hot rolling step, or is hot rolled upon
repeating afi~er momentary cooling, it is subjected to hot
rolling at a finishing temperature of Ar3 - 50°C to Ar3 +
120°C, cooled at an average cooling rate of more than
5°C/sec in a run-out table, and then coiled at a
temperature of no greater than 350°C; and
(10) a method of producing a dual-phase high-
strength hot-rolled steel sheet having high impact energy
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absorption properties according to (9) above,
characterized in that at the finishing temperature for
hot rolling in a range of Ar3 - 50°C~to Ar3 + 120°C, the
hot rolling _~s carried out so that the metallurgy
parameter A satisfies inequalities (1) and (2) below, the
subsequent average cooling rate in the run-out table is
at least 5°C/sec, and the coiling is accomplished so that
the relationship between the above-mentioned metallurgy
parameter A and the coiling temperature (CT) satisfies
inequality (:3) below.
9 <_ logA <_ 18 (1)
oT <_ 21 x logA - 61 (2)
CT <- 6 :.c logA + 242 ( 3 )
(11) The method of producing a dual-phase type high-
strength cold rolled steel sheet having high impact
energy abosorption properties according to the invention
is a method of producing a dual-phase type high-strength
cold rolled steel sheet with excellent dynamic
deformation properties according to (1) to (8) above,
characterized in that after a continuous cast slab is fed
directly from casting to a hot rolling step, or is hot
rolled upon :reheating after momentary cooling, it is hot
rolled, the :hot-rolled and subsequently coiled steel
sheet is cold-rolled after acid pickling, and during
annealing in a continuous annealing step for preparation
of the final product, it is heated to a temperature
between Aci and Acz and subjected to the annealing while
held in this temperature range for at least 10 seconds,
and then cooled at a cooling rate of more than 5°C/sec;
and
(12) a :method according to (11) above for producing
a dual-phase type high-strength cold rolled steel sheet
having high impact energy absorption properties according
to (1) to (8) above, characterized in that in the
continuous annealing step, the cold rolled steel sheet is
heated to a temperature (To) between Acl and Ac3 and
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subjected to the annealing while held in this temperature
range for at least 10 seconds, and for subsequent
cooling, it is cooled to a secondary~cooling start'
temperature (Tq) in the range of 550°C-To at a primary
cooling rate of 110°C/sec and then cooled to a secondary
cooling end temperature (Te) which is no higher than Tem
determined by the chemical compositions and annealing
temperature (To), at a secondary cooling rate of
10200°C/sec.
Brief Descrix~tion of the Drawinas
Fig. 1 is a graph showing the relationship between
the absorption energy (Eab) of a shaped member during
collision and the material strength (S), according to the
invention.
Fig. 2 _~s a perspective view of a shaped member for
measurement of impact absorption energy for Fig. 1.
Fig. 3 is a graph showing the relationship between
the work hardening coefficient and dynamic energy
absorption for a steel sheet.
Fig. 4 .is a graph showing the relationship between
the yield strength x work hardening coefficient and the
dynamic energy absorption for a steel sheet.
Fig. 5 .is a general view of a "hat model" used in
the impact crush test method relating to Figs. 3 and 4.
Fig. 6 is a cross-sectional view of the shape of the
test piece of Fig. 5.
Fig. 7 is a schematic view of the impact crush test
method relating to Figs. 3-6.
Fig. 8 is a graph showing the relationship between
TS and the difference between the average value 6dyn of
the deformation stress in the range of 310% of
equivalent strain when deformed in a strain rate range of
5 x lOZ - 5 x 103 (1/S) and TS, as an index of the impact
energy absorption property upon collision, according to
the invention.
Fig. 9 is a graph showing the change in the
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static/dynam.ic ratio with tempered rolling for an example
of the invention and a comparative example.
Fig. 10 is a graph showing the relationship between
oT and the m~atallurgy parameter A for a hot-rolling step
according to the invention.
Fig. 11 is a graph showing the relationship between
the coiling temperature and the metallurgy parameter A
for a hot-rolling step according to the invention.
Fig. 12 is a graph showing the annealing cycle for
continuous annealing according to the invention.
Best Mode fo:r Carrying Out the Invention
Impact .absorbing members such as front side members
of automobiles are produced by bending and press forming
of steel sheets. Because impacts during automobile
collisions are absorbed by such members which have
undergone press forming, they must have high impact
absorption properties even after having undergone the
pre-deformation corresponding to the press forming. At
the current time, however, no attempt has been made to
obtain high-strength steel sheets with excellent impact
absorption properties as actual members, with
consideration of both the increase in the deformation
stress by press forming and the increase in deformation
stress due to a higher strain rate, as was mentioned
above.
As a result of much experimentation and research
with the aim of achieving this purpose, the present
inventors have found that steel sheets with a dual-phase
(DP) structure are ideal as high-strength steel sheets
with excellent impact absorption properties for actual
members which are press formed as described above. It
was demonstrated that such steel sheets with a dual-phase
microstructure, which is a composite microstructure
wherein the dominating phase is a ferrite phase
responsible for the increase in deformation resistance by
an increased. strain rate, and the second phase includes a
hard martensite phase, have excellent dynamic deformation
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properties. That is, it was found that high dynamic
deformation properties are exhibited when the
microstructure of the final steel sheets is a composite
structure wherein the dominating phase is ferrite and
another low temperature product phase includes a hard
martensite phase at a volume fraction of 350% after
deformation at 5% equivalent strain of the steel sheet.
Concerning the volume fraction of 350% for the hard
martensite phase, since high-strength steel sheets and
even steel sheets with high dynamic deformation
properties cannot be obtained if the martensite phase is
less than 3%,. the volume fraction of the martensite phase
must be at least 3%. Also, if the martensite phase
exceeds 50%, this results in a smaller volume fraction of
the ferrite phase responsible for greater deformation
resistance due to increased deformation speed, making it
impossible to obtain steel sheets with excellent dynamic
deformation properties compared to static deformation
strength while also hindering press formability, and
therefore it was found that the volume fraction of the
martensite phase must be 350%.
The present inventors then pursued experimentation
and research based on these findings and, as a result,
found that a:Lthough the degree of pre-deformation
corresponding to press forming of impact absorbing
members such as front side members sametimes reaches a
maximum of over 20%, depending on the location, the
majority are locations with 0%~10% of equivalent strain,
and that by understanding the effect of pre-deformation
in this range, it is possible to estimate the behavior of
the member a5 a whole after pre-deformation.
Consequently, according to the invention, a deformation
of from 0% to 10% of equivalent strain was selected as
the amount of pre-deformation applied to members during
press forming.
Fig. 1 is a graph showing the relationship between
the absorption energy (Eab) of a press formed member
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during collision and the material strength (S), for the
different st~ael types shown in Table 5, according to an
example to bra described later. The material strength S
is the tensile strength (TS) according to the common
tensile test. The member absorption energy (Eab) is the
absorption energy in the lengthwise direction (direction
of the arrow) along a press formed member such as shown
in Fig. 2, upon collision with a 400 kg mass weight at a
speed of 15 m/sec, to a crushing degree of 100 mm. The
shaped member in Fig. 2 consists of a 2.0 mm-thick steel
sheet formed into a hat-shaped section 1 with a steel
sheet 2 of the same thickness and the same type of steel,
joined together by spot welding, the hat-shaped section 1
having a corner radius of 2 mm, and with spot welding
points indicated by 3.
From Fig. 1 it is seen that the member absorption
energy (Eab) tends to increase with the strength of
materials under normal tensile testing, though with
considerable variation. Here, the materials in Fig. 1
were subjected to pre-deformation of more than 0% and
less than or equal to 10% of equivalent strain, and then
the static deformation strength 6s when deformed in a
strain rate range of 5 x 10 4 - 5 x 10 3 ( s 1 ) and the
dynamic deformation strength ad when deformed in a strain
rate range of 5 x lOZ - 5 x 103 ( s 1 ) after the pre-
deformation, were measured. As a result, a
classification was possible based on (~d - mss). The
symbols plotted in Fig. 1 were as follows:
o: (ad - 6s) < 60 MPa with any pre-deformation of
more than 0% and less than or equal to 10%;
~: 60 M.Pa <_ (6d - as) with any :pre-deformation in
the above range, and 60 MPa _< (6d - as) < 80 MPa
with pre-deformation of 5%;
1: 60 M.Pa <_ (ad - cs) with any pre-deformation in
the above range, and 80 MPa <_ (~d - os) < 100 MPa
with. pre-deformation of 5%;
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~: 60 MF~a -< (ad - os) with any pre-deformation in
the above range, and 100 MPa <- (od - 6s) with
pre-deformation of 5%.
Also, when 60 MPa <_ (ad - 6s) with any pre-
y deformation i.n the range of more than 0% and less than or
equal to 10% of equivalent strain, the values for member
absorption energy (Eab) during collision was equal to or
greater than the values predicted from the material
strength S, thus indicating steel sheets with excellent
dynamic deformation properties as impact absorbing
members for collision. These predicted values are those
shown in the curve in Fig. 1, represented by Eab =
0.062S°'8. Consequently, (od - os) must be at least 60
MPa.
For improved impact absorption properties, it is
basically important to increase the work hardening
coefficient, specifically to at least 0.13, and
preferably at: least 0.16; by controlling the yield
strength and the work hardening coefficient to specified
ranges it is possible to achieve excellent impact
absorption pi_-operties, and for improved press formability
it is effective to design the volume percentage and
particle size of the martensite to within a specified
range.
Fig. 3 shows the relationship between the work
hardening coefficient of a steel sheet and the dynamic
energy absorption which indicates the member impact
absorption properties, for a class of: materials with the
same yield strength. Here it is shown that increased
work hardening coefficients of the steel sheets result in
improved member impact absorption properties (dynamic
energy absorption), and that the work hardening
coefficient of a steel sheet can properly indicate the
member impact absorption properties so long as the yield
strength class is the same. Also, when the yield
strengths differ, as shown in Fig. 4, the yield strength
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x work hardening coefficient can be an indicator of the
member impact absorption properties. While the work
hardening coefficient was expressed in terms of an n
value of 5~~:10~ strain in consideration of the strain
undergone by members during press forming, from the
viewpoint of improving the dynamic energy absorption,
work hardening coefficients of under 5~ strain or work
hardening coE~fficients of even more than 10~ strain may
be preferred.
The dynamic energy absorptions for members shown in
Fig. 3 and F:ig. 4 were determined in the following
manner. Specifically, the steel sheet was shaped into
the member shape shown in Fig. 6 (corner R = 5 mm) and
spot welded at 35 mm pitch using an electrode with a tip
radius of 5..'i mm at a current of 0.9 times the expulsion
current, and then after baking and painting treatment at
170°C x 20 minutes, an approximately 150 Kg falling
weight was dropped from a height of about 10 m to crush
the member in its lengthwise direction, and the
displacement work where displacement = 0-150 mm is
calculated from the area of the corresponding load
displacement diagram to determine the dynamic energy
absorption. A schematic illustration of this test method
is shown in :Eig. 7. In Fig. 5, 4 is a worktop, 5 is a
test piece a:nd 6 is a spot welding section.
In Fig. 6, 7 is a hat-shaped test piece and 8 is a
spot welding section. In Fig. 7, 9 is a worktop, 10 is a
test piece, 11 is a falling weight (150 kg), 12 is a
frame, and 13 is a shock absorber. The work hardening
coefficient and yield strength of each steel sheet was
determined in the following manner. The steel sheet was
shaped into a JIS-#5 test piece (gauge length: 50 mm,
parallel width: 25 mm), subjected to tensile test at a
strain rate of 0.001 (s1) to determine the yield
strength and work hardening coefficient (n value at
5%~10% strain). The steel sheet used had a sheet
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thickness of 1.2 mm and the steel sheet composition
contained C at 0.020.25 wt~, either or both Mn and Cr at
a total of 0.153.5 wt% and one or more of Si, A1 and P
at a total o.f 0.024.0 wt%, with the remainder Fe as the
main component .
Fig. 8 .is a graph showing the relationship between
the average value adyn of the deformation stress in the
range of 310% of equivalent strain when deformed in a
strain rate .range of 5 x 102 - 5 x 103 ( s 1 ) and the
static material strength (TS), as an index of the impact
energy absorption property upon collision according to
the invention, where the static material strength (TS) is
the tensile strength (TS: MPa) in the static tensile test
as measured .in a strain rate range of 5 x 104 - 5 x 103
( s-1 ) .
As mentioned above, impact absorbing members such as
front side members have a hat-shaped cross-sectional
shape, and a;s a result of analysis of deformation of such
members upon crushing by high-speed collision, the
present inventors have found that despite deformation
proceeding up to a high maximum strain of over 40%, at
least 70% of the total absorption energy is absorbed in a
strain range of 10% or lower in a high-speed stress-
strain diagram. Therefore, the dynamic deformation
resistance with high-speed deformation at 10% or lower
was used as the index of the high-speed collision energy
absorption property. In particular, since the amount of
strain in the range of 310% is most important, the index
used for the impact energy absorption property was the
average stress: adyn in the range of 310% of equivalent
strain when deformed in a strain rate range of 5 x 10z -
5 x 103 (s1) high-speed tensile deformation.
The average stress: edyn of 310% upon high-speed
deformation generally increases with increasing static
tensile strength {maximum stress (TS: MPa) in a static
tensile test measured in a stress rate range of 5 x
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4 - 5 x 10 ' ( s-1 ) } of the steel material prior to pre-
deformation or baking treatment. Consequently,
increasing the static tensile strength (which is
synonymous w:Lth the static material strength) of the
5 steel material directly contributes to an improved impact
energy absorption property of the member. However,
increased strength of the steel results in poorer press
formability :into members, making it difficult to obtain
members with the necessary shapes. Consequently, steels
10 having a high 6dyn with the same tensile strength TS are
preferred. :It was found that, based on this
relationship, steel sheets wherein the average value cdyn
(MPa) of the deformation stress in the range of 3~10~ of
equivalent strain when deformed in a strain rate range of
5 x 102 - 5 ~: 103 ( s 1 ) , after pre-deformation of more
than 0~ and .Less than or equal to 10~ of equivalent
strain satisfies the inequality: adyn ? 0.766 x TS + 250
as expressed in terms of the tensile strength (TS: MPa)
in the static tensile test as measured in a strain rate
range of 5 x 10 4 - 5 x 10 3 ( s-1 ) prior to pre-
deformation, have higher impact energy absorption
properties as actual members compared to other steels,
and that the impact energy absorption property is
improved without increasing the overall weight of the
member, making it possible to provide high-strength steel
sheets with high dynamic deformation resistance.
Also, although the details are still unclear, it has
been discovered that steel sheets with excellent dynamic
deformation properties can be obtained when, as shown in
Fig. 9, YS(0)/TS'(5) is no greater than 0.7, which amount
is dependent on the initial microstructure, the amount of
solid solution elements in the low temperature product
phase other than the martensite phase and the main
ferrite phase, and the deposited state of carbides,
nitrides and carbonitrides. Here, YS(0) is the yield
strength, and TS'(5) i_s the tensile strength (TS') in the
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static tensile test with pre-deformation at 5% of
equivalent si~rain or after further bake hardening
treatment (BH treatment). It was also demonstrated that
steel sheets with even more excellent dynamic deformation
properties can be obtained when the yield strength: YS(0)
x work hardening coefficient is at least 70.
Furthermore, it is known that dynamic deformation
strength is usually expressed in the form of the power of
the static tensile strength, and as the static tensile
strength inc:ceases, the difference between the dynamic
deformation atrength and the static deformation strength
decreases. l3owever, a small difference between the
dynamic deformation strength and the static deformation
strength will mean that no greater improvement in the
impact absorption properties can be expected. From this
standpoint, it is preferred for the value of (od - os) to
be in a range which satisfies the following inequality,
( 6d - as ) ? 4 . 1 x os°'a - as .
The microstructure of a steel sheet according to the
invention will now be described in detail. As already
mentioned, the martensite is at a volume fraction of
350%, and preferably 330%. The average grain size of
the martensite is preferably no greater than 5 Vim, and
the average grain size of the ferrite is preferably no
greater than 10 Vim. That is, the martensite is hard, and
contributes to a decrease in the yield ratio and an
improvement in the work hardening coefficient, by
producing a mobile dislocations primarily in adjacent
ferrite grains; however, by satisfying the restrictions
mentioned above it is possible to disperse fine
martensite in the steel, so that the improvement in the
properties spreads throughout the entire steel sheet. In
addition, this dispersion of fine martensite in the steel
can help to avoid deterioration in the hole expansion
ratio and tensile strength x total elongation, which is
an adverse effect of the hard martensite. Also, because
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it is possible to reliably achieve work hardening
coefficient ._ 0.130, tensile strength x total elongation
18,000 and hole expansion ratio >- 1.2, it is thereby'
possible to improve the impact absorption properties and
press formability.
With a martensite volume fraction of less than 3%,
the yield ratio becomes larger while the press formed
member cannot: exhibit an excellent work hardening
property (work hardening coefficient >- 0.130) after it
has undergone collision deformation, and since the
deformation resistance (load) stays at a low level, and
the dynamic energy absorption is low preventing
improvement in the impact absorption properties. On the
other hand, with a martensite volume fraction of greater
than 50%, the yield ratio becomes larger while work
hardening coefficient is reduced, and deterioration also
occurs in the tensile strength x total elongation and the
hole expansion ratio. From the standpoint of press
formability, the volume fraction of the martensite is
preferred to be no greater than 30%.
Also, the ferrite is present at a volume fraction of
preferably at; least 50%, and more preferably at least
70%, and its average grain size (mean circle equivalent
diameter) is preferably no greater than 10 Vim, and more
preferably no greater than 5 Vim, with the martensite
preferably adjacent to the ferrite. This aids the fine
dispersion oj= the martensite in the ferrite matrix, while
effectively extending the property-improving effect,
beyond simply a local effect, to the entire steel sheet,
favorably acting to prevent the adverse effects of the
martensite. The structure of the remainder present with
the martensite and ferrite may be a mixed structure
comprising a combination of one or. more from among
pearlite, ba:inite, retained 'y, etc., and although
primarily ba:inite is preferred in cases which require
hole expansion properties, since retained y undergoes
work-induced transformation into martensite by press
CA 02283924 1999-09-13
- 18 -
forming, experimental results have shown that including
retained aust:enite prior to press forming has an-effect
even in preferred small amounts (5% or less).
Also, from the standpoint of impact absorption
properties and press formability it is preferred for the
ratio of the martensite and ferrite particle sizes to be
no greater than 0.6, and the ratia of the hardnesses to
be at least .1.5.
The rest:rictions on the values for the chemical
components of-_ dual-phase type high-strength steel sheets
with excellent dynamic deformation properties according
to the inveni:ion, and the reasons for those restrictions,
will now be Explained.
Dual-phase type high-strength steel sheets with
excellent dynamic deformation properties which are used
according to the invention are steel sheets containing
the following chemical compositions, in terms of weight
percentage: c~ at 0.02~0.25%, either or both Mn and Cr at
a total of 0.15~3.5%, one or more from among Si, Al and P
at a total o:E 0.02~4.0%, if necessary also one or more
from among N:i, Cu and Mo at a total of no more than 3.5%,
one or more :from among Nb, Ti and V at no more than
0.30%, and either or both Ca and REM at 0.0005~0.01% for
Ca and 0.0050.05% for REM, with the remainder Fe as the
primary component. They are also dual-phase type high
strength ste~=1 sheets with excellent dynamic deformation
properties which contain, if necessary, one or more from
among B (<_0.01), S (__<0.01%) and N (<_0.02%). These
chemical components and their contents (percent by
weight) will now be discussed.
C: C is the element which most strongly affects the
microstructure of the steel sheet, and if its content is
too low it will become difficult to obtain martensite
with the desired amount and strength. Addition in too
great an amount leads to unwanted carbide precipitation,
inhibited increase in deformation resistance at higher
strain rates and overly high strength, as well as poor
CA 02283924 1999-09-13
- 19 -
press formab_Llity and weldability; the content is
therefore 0.020.25 wt~.
Mn, Cr: Mn and Cr have an effect of stabilizing
austenite and guaranteeing sufficient martensite, and are
also solid solution hardening elements; they must
therefore be added in a minimum amount of 0.15 wt~, but
if added in i~oo much the aforementioned effect becomes
saturated thus producing adverse effects such as
preventing ferrite transformation, and thus they are
added in the maximum amount of 3.5 wt~.
Si, A1, P: Si and A1 are useful elements for
producing martensite, and they promote production of
ferrite and suppress precipitation of carbides, thus
having the e:Efect of guaranteeing sufficient martensite,
as well as a solid solution hardening effect and a
deoxidization effect. P can also promote martensite
formation and solid solution hardening, similar to A1 and
Si. From this standpoint, the minimum amount of Si + A1
+ p added muat be at least 0.02 wt~. On the other hand,
excessive addition will saturate this effect and result
instead in b:rittleness, and therefore the maximum amount
of addition is no more than 4.0 wt~. In particular, when
an excellent surface condition is required, Si scales can
be avoided b:y adding Si at no greater than 0.1 wt%, and
conversely b:y adding it at 1.0 wto or greater Si scales
can be produ~~ed over the entire surface so that they are
not conspicuous. Also, when excellent secondary
workability, toughness, spot weldability and recycling
properties are required, the P content may be kept at no
greater than 0.05%, and preferably no greater than 0.02%.
Ni, Cu, Mo: These elements are added when necessary,
and are austenite-stabilizing elements similar to Mn,
which increase the hardenability of the steel, and are
effective for adjustment of the strength. From the
standpoint of weldability and chemical treatment, they
can be used when the amounts of C, Si, A1 and Mn are
restricted, but if the total amount of these elements
CA 02283924 1999-09-13
- 20 -
added exceeds 3.5 wt% the dominant ferrite phase will
tend to be hardened, thus inhibiting the increase in
deformation .resistance by a greater strain rate, as well
as raising the cost of the steel sheet; the amount of
these elements added is therefore 3.50 wt% or lower.
Nb, Ti, V: These elements are added when necessary,
and are effe~~tive for strengthening the steel sheet
through formation of carbides, nitrides and
carbonitridea. However, when added at greater than 0.3
wt% they are deposited in large amounts in the dominant
ferrite phase=_ or at the grain boundaries as carbides,
nitrides and carbonitrides, becoming a source of the
mobile dislocation during high speed deformation, and
inhibiting t:he increase in deformation resistance by
greater strain rates. In addition, the deformation
resistance of the dominant phase becomes higher than
necessary, thus wasting the C and leading to higher
costs; the maximum amount to be added is therefore 0.3
wt%.
B: B is an element which is effective for
strengthening since it improves the hardenability of the
steel by suppressing production of ferrite, but if it is
added at greater than 0.01 wt% its effect will be
saturated, and therefore B is added at a maximum of 0.01
wt%.
Ca, REM: Ca is added to at least 0.0005 wt% for
improved press formability (especially hole expansion
ratio) by shape control (spheroidizat~ion) of sulfide-
based inclusions, and the maximum amount thereof to be
added is 0.01 wt% in consideration of effect saturation
and the adverse effect due to increase in the
aforemention?d inclusions (reduced hole expansion ratio).
For the same reasons, REM is added in an amount of from
0.005% to 0.05 wt%.
S: The amount of S is no greater than 0.01 wt%, and
preferably no greater than 0.003 wt%, from the standpoint
of press formability (especially hole expansion ratio) by
CA 02283924 1999-09-13
- 21 -
sulfide-based inclusions, and reduced spot weldability.
The method of applying the pre-deformation according
to the inveni~ion will now be explained. The pre-
deformation may be press forming for member shaping, or
it may be working with a tempering rolling or tension
leveler which applied to the steel sheet material prior
to its press forming. In this case, either or both a
tempering roller and tension leveler may be used. That
is, the mean, used may include a tempering rolling, a
tension leve:Ler, or a tempering roller and tension
leveler. The steel sheet material may also be subjected
to press forming after being worked with a tempering
rolling or tension leveler. The amount of pre-
deformation applied with the tempering rolling and/or
tension leve:Ler, i.e. the degree of plastic deformation
(T), will differ depending on the initial dislocation
density, and T should be small if the initial density is
large. Also, with few solid solution elements the
introduced dislocations cannot be fixed, and high dynamic
deformation :properties cannot be guaranteed.
Consequently, it was found that the plastic deformation
(T) is determined based on the ratio between the yield
strength YS(0) and the tensile strength TS'(5) in the
static tensile test with pre-deformation at 50 of
equivalent strain or after further bake hardening
treatment (BH treatment), or YS(0)/TS'(5). That is,
YS(0)/TS'(5) is an indicator of the sum of the initial
dislocation density and the dislocation density
introduced by 5% deformation, and the amount of the solid
solution elements; it may be concluded that a smaller
YS(0)/TS'(5) means a higher initial dislocation density
and more of the solid solution elements. YS(0)/TS'(5) is
therefore no greater than 0.7, and is preferably provided
according to the following equation:
2.5 ~YS(0)/TS'(5) - 0.5} + 15 ? T >_ 2.5
~YS(0)/TS'(5) - 0.5} + 0.5
wherein the upper limit for T is determined from the
CA 02283924 1999-09-13
- 22 -
standpoint o:E press formability including impact
absorption property and flexibility.
A method of producing a dual-phase type high
strength hot rolled steel sheet and a cold rolled steel
sheet with e:~ccellent dynamic deformation properties
according to the invention will now be explained. In
this product:ion method, a continuous cast slab is fed
directly from casting to a hot rolling step, or is hot
rolled upon ceheating after momentary, cooling. Thin
gauge continuous casting and continuous hot rolling
techniques (endless hot rolling) may be applied for the
hot rolling :in addition to normal continuous casting, but
in order to avoid a lower ferrite volume fraction and a
coarser average grain size of the thin steel sheet
microstructure, the bar (cast strip) thickness at the hot
rolling approach side (the initial steel bar thickness)
is preferred to be at least 25 mm. At less than 25 mm,
the mean circle equivalent size of ferrite of the steel
sheet is made=_ coarser, while it is also a disadvantage
against obtaining the desired martensite. The final pass
rolling speed for the hot rolling is preferred to be at
least 500 mpm and more preferably at least 600 mpm, in
light of the problems described above. At less than 500
mpm, the mean circle equivalent diameter of ferrite of
the steel sheet is made coarser, while it is also a
disadvantage against obtaining the desired martensite.
The finishing temperature for the hot rolling is
from Ar3 - 50°C to Ar3 + 120°C. At lower than Ar3 - 50°C,
deformed ferrite is produced, with inferior work
hardening property and press formability. At higher than
Ar3 + 120°C, and the mean circle equivalent size of
ferrite of the steel sheet is made coarser, while it is
also becomes difficult to obtain the desired martensite.
The average cooling rate for cooling in the run-out
table is at least 5°C/sec. At less than 5°C/sec it
becomes difficult to obtain the desired martensite.
CA 02283924 1999-09-13
- 23 -
The coiling temperature is no higher than 350°C. At
higher than :350°C it becomes difficult to obtain the
desired martensite.
According to the invention, it was found
particularly that a correlation exists between the
finishing temperature in the hot rolling step, the
finishing approach temperature and the coiling
temperature. That is, as shown in Fig. 10 and Fig. 11,
specific conditions exist which are determined primarily
between the :Finishing temperature, finishing approach
temperature and the coiling temperature. Specifically,
the hot rolling is carried out so that when the finishing
temperature :For hot rolling is in the range of Ar3 - 50°C
to Ar3 + 120°C, the metallurgy parameter A satisfies
, inequalities (1) and (2). The above-mentioned metallurgy
parameter A may be expressed by the following equation.
A = Ef~ x: exp{ ( 75282 - 42745 x Cep) / [ 1 . 978 x (FT +
273)]}
where FT: finishing temperature (°C)
Ceq: carbon equivalents = C + Mneq/6 ( % )
MnE,q: manganese equivalents = Mn + (Ni + Cr +
Cu + Mo)/2 (%)
Ef: final pass strain rate (s 1)
(v/JRxhl) x (1/Jr) x In X1/(1-r)}
hl: final pass approach sheet thickness
hz: final pass exit sheet thickness
r . (hl - hz) /hl
R . roll radius
v . final pass exit speed
0T: fin.ishing temperature (finishing final pass exit
temperature) - finishing approach temperature
(finish.ing first pass approach temperature)
Ar3: 901. - 325 C% + 33 Si % - 92 Mneq
Thereafter, it is preferred for the average cooling
rate on the :run-out table to be at least 5°C/sec, and the
coiling to b~a carried out under conditions such that the
CA 02283924 1999-09-13
- 24 -
relationship between the metallurgy parameter A and the
coiling temperature (CT) satisfies inequality (3).
9 <_ logA <- 18 ~ (1)
eT <- 21 x logA - 61 (2)
CT <_ 6 ;K logA + 242 (3)
In inequality (1) above, a log A of less than 9 is
unacceptable from the viewpoint of production of retained
martensite and refinement of the microstructure, while it
will also reault in an inferior dynamic deformation
resistance adyn and 5~10~ work hardening property. Also,
if log A is -to be greater than 18, massive equipment will
be required to achieve it. With inequality (2), if the
condition of inequality (2) is not satisfied it will be
impossible to obtain the desired martensite, and the
dynamic deformation resistance ~dyn and 5~10~ work
hardening property, etc. will be inferior. The lower
limit for eT is more flexible with a lower log A as
indicated by inequality (2). Furthermore, if the
relationship with the coiling temperature in inequality
(3) is not satisfied, there will be an adverse effect on
ensuring the amount of martensite, while the retained y
will be excessively stable even if retained y can be
obtained, it will be impossible to obtain the desired
martensite during deformation, and the dynamic
deformation resistance 6dyn and 5~10o work hardening
property, etc. will be inferior. The limit for the
coiling temperature is more flexible with a higher log A.
The cold rolled sheet according to the invention is
then subjected to the different steps following hot-
rolling and coiling and is cold rolled and subjected to
annealing. The annealing is ideally continuous annealing
through an annealing cycle such as shown in Fig. 12, and
during the annealing of the continuous annealing step, it
must be kept. for at least 10 seconds in the temperature
range of Acl - Ac3. At less than Acl austenite will not
be produced and it will therefore be impossible to obtain
CA 02283924 1999-09-13
- 25 -
martensite thereafter, while at greater than Ac3 the
austenite mon.ophase structure will be coarse, and it will
therefore be impossible to obtain the desired average
grain size for the martensite. Also, at less than 10
seconds the a.ustenite production will be insufficient,
making it impossible to obtain the desired martensite
thereafter. The maximum residence time is preferably no
greater than 200 seconds, from the standpoint of avoiding
addition to the equipment and coarsening of the
microstructure. The cooling after this annealing must be
at an average cooling rate of at least 5°C/sec. At less
than 5°C/sec the desired space factor for the martensite
cannot be achieved. Although there is no particular
upper limit here, it is preferably 300°C/sec when
considering temperature control during the cooling.
According to the invention, the cooled steel sheet
is heated to a temperature To from Acl- Ac3 in the
continuous annealing cycle shown in Fig. 12, and cooled
under cooling conditions provided by a method wherein
cooling to a secondary cooling start temperature Tq in
the range of 550°C-To at the primary cooling rate of
110°C/sec i=~ followed by cooling to a secondary cooling
end temperature Te which is no higher than a temperature
Tem which is determined by the chemical compositions of
the steel anti annealing temperature To, at a secondary
cooling rate of 10200°C/sec. This is a method whereby
the cooling end temperature Te in the continuous
annealing cycle shown in Fig. 12 is represented as a
function of t:he chemical compositions and annealing
temperature, and is kept under a given critical value.
After cooling to Te, the temperature is preferably held
in a range of Te - 50°C to 400°C for up to 20 minutes
prior to cooling to room temperature.
Here, Tem is the martensite transformation start
temperature nor the retained austenite at the quenching
start point ~Cq. That is, Tem is defined by Tem = T1 -
CA 02283924 1999-09-13
- 26 -
T2, or the difference between the value excluding the
effect of the C concentration in the austenite (T1) and
the value indicating the effect of the C concentration
(T2). Here, T1 is the temperature calculated from the
solid solution element concentration excluding C, and T2
is the temperature calculated from the C concentration in
the retained austenite at Acl and Ac3 determined by the
chemical compositions of the steel and Tq determined by
.,
the annealing temperature To. Ceq represents the carbon
equivalents in the retained austenite at the annealing
temperature To. Thus, T1 is expressed as:
T1 = 561 - 33 x {Mn% + (Ni + Cr + Cu + Mo)/2}
and T2 is expressed in terms of:
Acl = 72:3 - 0.7 x Mn% - 16.9 x Ni% + 29.1 x Si% +
16.9 x Cr%,
Ac3 = 911) - 203 x (C% ) l~z - 15 . 2 x Ni% + 44 . 7 x Si% +
104 x V% + 31.5 x Mo% - 30 x Mn% - 11 x Cr% - 20 x
Cu% + 70 x P% + 40 x A1% + 400 x Ti%,
and the annealing temperature To, and when
Ceq'~ - (Ac3 - Acl) x C/(To - Acl) + (Mn + Si/4 + Ni/7
+ Cr + Cu + 1.5 Mo)/6
is greater than 0.6, T2 - 474 x (Ac3 - Acl) x C/(To -
Acl) ,
and when it is 0.6 or less, T2 - 474 x (Ac3 - Acl) x C/~3
x (Ac3 - Ac1) x C + [(Mn + Si/4 + Ni/7 + Cr + Cu + 1.5
Mo)/2 - 0.85)] x (To - Aci).
In other words, when Te is equal to or greater than
Tem, the desired martensite cannot be obtained. Also, if
Toa is 400°C or higher, the martensite obtained by
cooling is tempered, making it impossible to achieve
satisfactory dynamic properties and press formability.
On the other hand, if Toa is less than Te - 50°C,
additional ccoling equipment is necessary, and greater
variation will result in the material due to the
difference between the temperature of the continuous
annealing furnace and the temperature of the steel sheet;
CA 02283924 1999-09-13
- 27 -
this temperature was therefore determined as the lower
limit. Also, the upper limit for the holding time was
determined to be 20 minutes, because'when it is longer
than 20 minutes it becomes necessary to expand the
equipment.
By employing the chemical composition and production
method described above, it is possible to produce a dual-
phase type high-strength steel sheet with excellent
dynamic deformation properties, wherein the
microstructure of the steel sheet is a composite
microstructure wherein the dominating phase is ferrite,
and the second phase is another low temperature product
phase containing martensite at a volume fraction from
3%~50% after shaping and working at 5% equivalent strain,
and wherein t:he difference between the quasi-static
deformation ~~trength cs when deformed in a strain rate
range of 5 x 104 - 5 x 103 (1/s) after pre-deformation
of more than 0% and less than or equal to 10% of
equivalent strain, and the dynamic deformation strength
~d measured ~n a strain rate range of 5 x 102 - 5 x 10'
(1/s) after the aforementioned pre-deformation, i.e.
(6d - as), is at least 60 MPa, and the work hardening
coefficient at 510% strain is at least 0.13. The steel
sheets according to the invention may be made into any
desired product by annealing, tempering rolling,
electronic coating or hot-dip coating.
Examples
The present invention will now be explained by way
of examples.
(Example 1)
The 26 steel materials listed in Table 1 (steel nos.
126) were heated to 10501250°C and subjected to hot
rolling, cooling and coiling under the production
conditions listed in Table 2, to produce hot rolled steel
sheets. As shown in Table 3, the steel sheets satisfying
the chemical composition conditions and production
CA 02283924 1999-09-13
- 28 -
conditions according to the invention have a dual-phase
structure with a martensite volume fraction of at least
3~ and no greater than 50~, and as shown in Fig. 4, the
mechanical properties of the hot rolled steel sheets
indicated excellent impact absorption properties as
represented by a work hardening coefficient of at least
0.13 at 5~10~: strain, ~d - os >_ 60 MPa, and odyn ? 0.766
x TS + 250, while also having suitable press formability
and weldabili.ty.
CA 02283924 1999-09-13
- 29 -
cr m ~no .-io 0 0 0 ~~o o m n ~no ~mn ~ o o ~ o ,mn ~ o
M O M v0r1N .-IO 00N ~ N O O~.-IO~.w1N 00 0000rin r1
N N N
r-1r1r1~--Ir-I.-Irir1,-ir-i.-I.-Iri.-1r-1r1.-ie-Ir-1.-I.-i.-1O r-I.-1O
OVf~N O v1W700~.OvY O O~N O~00OWO N vY'I~~l1f~~Ya0t~N
O" N N N M M M N N ~:VN ~ N M N .-iM N d N N M M N N M
O
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U M M M O M ~De-IN ~-iN ~ M ~1M O N O~O r~N N GO00~-iI~~-1
r-Irir-Ir1r-1r-ir1r-i~-ir-ir-If-1r1.-Iflr1O N r1.-IN r-iO .-i'-IO
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n N 00-I N N N N .-IN N e-IN rie-~Ia0a~e-Ir-i
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O O O O O O O O O 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0
0
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O O O O O O O O O O O O r-IO O ~D O r-i~flr1O
O O O e-IO N O N 016~riN M OJ00r1n
N N M O M ~Or--1N e-ICO CON ~Y
~ -io ~ r .-i~ ~ o ~ ~ o
U r-i~ ~ ~ ~ ri~ ~ ~ o ~ ,~~ ~ . r
O -iN O l r-ir1r--tO O v0O O O N O O r-1N
'-Ir1O r-1r-Ir1O c0. O O r O O O O~Q~r1N O~O O N N ~D
' O ~ O O O
.-i O O N ~tO c0
1 O O -I-Ir-ir-1O O O N O O e-1O O O r-I
-Ir-1r-I~-1O O O O r ~ O
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o00om 0 0 0 ~ 0 0 0 0 0 0 0 0 0 0 0 ~ o 0
V o 0 0 ~ 0 0 0
0 0 0
0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0
a~
t'D n ~ OvO r-iN M d u1tDf~COOvO r-IN M ~ W D O
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CA 02283924 1999-09-13
- 30 -
Table 1 (cont.) Chemical compositions of steels
Ste~slTransformation Type
No. temperature
C
Acl Ac3 Ar3
1. 741 863 793 present invention
2 741 863 793 present invention
3 744 880 805 present invention
9: 756 871 809 present invention
5 709 863 756 present invention
E 706 816 731 present invention
7 726 851 794 present invention
8 733 874 791 present invention
712 834 774 present invention
1C1 722 830 787 present invention
17. 733 839 736 present invention
12 741 863 793 present invention
la 739 857 775 comparative
example
14E 741 863 793 comparative
example
1-'i 713 861 806 comparative
example
1f> 728 839 732 present: invention
17 740 887 802 present invention
18 767 889 763 present invention
19 735 870 807 present invention
2t) 736 862 798 present invention
2:1 753 860 751 present invention
2:? 704 810 713 present invention
2:3 720 837 801 present invention
24 717 82.6773 present invention
25 752 923 771 present invention
26 722 779 762 comparative
example
CA 02283924 1999-09-13
- 31 -
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o U l
U H G1 O O O r-I1 -I1 N N N N N N N M
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CA 02283924 1999-09-13
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(Example 2)
The 22 steel materials listed in Table 5 (steel nos.
2748) were heated to 10501250°C and subjected to hot
rolling, coo:Ling and coiling, followed by acid pickling
and then cold rolling under the conditions listed in
Table 6 to p=roduce cold rolled steel sheets.
Temperatures Acl and Ac3 were then calculated from the
chemical compositions for each steel, and the sheets were
subjected to heating, cooling and holding under the
annealing conditions listed in Table 6, prior to cooling
to room temperature. As shown in Table 7, the steel
sheets satisfying the chemical composition conditions and
production conditions according to the invention have a
dual-phase structure with a martensite volume fraction of
at least 3% and no greater than 50% and, as shown in Fig.
8, the mechanical properties of the hot-rolled steel
sheets indicated excellent impact absorption properties
as represented by a work hardening coefficient of at
least 0.13 at 510% strain, od - 6s :? 60 MPa, and adyn
0.766 x TS + 250, while also having suitable press
formability and weldability.
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The microstructure was evaluated by the following
method.
Identification of the ferrite, bainite, martensite
and residual structure, observation of the location and
measurement of the average grain size (mean circle
equivalent.di.ameter) was accomplished using a 1000
magnification optical micrograph with the thin steel
sheet rolling direction cross-section etched with a nital
and the reagent disclosed in Japanese Unexamined Patent
Publication rlo. 59-219473.
The proF>erties were evaluated by the following
methods.
A tensile test was conducted according to JISS
(gauge mark distance: 50 mm, parallel part width: 25 mm)
with a strain rate of 0.001/s and, upon determining the
tensile strength (TS), yield strength (YS), total
elongation (T. E1) and work hardening coefficient (n
value for 1%~~5% strain), the YS x work hardening
coefficient and TS x T. E1. were calculated.
The stretch flanging property was measured by
expanding a 20 mm punched hole from the burrless side
with a 30° cone punch, and determining the hole expansion
ratio (d/do) between the hole diameter (d) at the moment
at which the crack penetrated the plate thickness and the
original hollow diameter (do, 20 mm).
The spot. weldability was judged to be unsuitable if
a spot welding test piece bonded at a current of 0.9
times the expulsion current using an electrode with a tip
radius of 5 times the square root of the steel sheet
thickness underwent peel fracture when ruptured with a
chisel.
Industrial At?plicability
As explained above, the present invention makes it
possible to provide, in an economical and stable manner,
high-strengt'.z hot rolled steel sheets and cold rolled
steel sheets for automobiles which provide previously
CA 02283924 1999-09-13
- 42 -
unobtainable excellent impact absorption properties and
press formability and thus offers a markedly wider range
of objects and conditions for uses of high-strength steel
sheets.
