Sabih-Nemes

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Journal of Materials Processing Technology 209 (2009) 4292–4311
Contents lists available at ScienceDirect
Journal of Materials Processing Technology
journa l homepage: www.e lsev ier .com/ locate / jmatprotec
nternal ductile failure mechanisms in steel cold heading process
. Sabiha,∗, J.A. Nemesb
Department of Mechanical Engineering, McGill University, 817 Sherbrooke Ave West, Montreal, Quebec, Canada H3A 2K6
Engineering Division, Pennsylvania State University-Great Valley, 30 East Swedesford Road, Malvern, PA 19355-1443, United States
r t i c l e i n f o
rticle history:
eceived 22 February 2008
eceived in revised form 5 November 2008
ccepted 7 November 2008
eywords:
uctile failure
diabatic shear band
old heading
oid nucleation
nclusions
a b s t r a c t
The occurrence of internal ductile failure in cold-headed products presents a major obstacle in the fast
expanding cold heading (CH) industry. This internal failure may lead to catastrophic brittle fracture under
tensile loads despite the ductile nature of the material. Comprehensive testing and investigation method-
ologies were used to this work to reveal the complicated interplay of process and material parameters
contributing in the initiation and propagation of internal ductile failure in six CH quality AISI steel grades.
The metallurgical and microscopic investigations showed that internal ductile failure occurs pro-
gressively by void nucleation and growth mechanisms with increasing plastic strain inside the highly
localized adiabatic shear bands (ASBs). The void nucleation occurs by decohesion at second-phase par-
ticles, inclusion–matrix interfaces, grain boundaries and by particle or inclusion cracking. Therefore,
the number and morphology of any inclusions and second-phase particles are key factors in material
formability.
The metallurgical investigations showed that under compressive loading conditions, the nature of the
metal flow pattern promotes different rates of material flow around the inclusions and stringers which
supports decohesion and void nucleation since the early stages of deformation. At advanced stages of
deformation, the metal flow pattern contributes to the ASB localization in supporting void growth and
coalescence along the band leading to narrow void sheets.
All tested materials in this work experienced ductile failure by void nucleation and coalescence, forming
cracks along the ASBs. The ductile failure of each material was the result of the contribution of all the
mechanisms of void nucleation at the inclusion–matrix interface, second phase–matrix interface and at
the grain boundaries. However, the level of contribution of each mechanism in the final ductile failure
erial
varied depending on mat
. Introduction
The occurrence of any failure is a major limitation governing the
imits of any forming process. Therefore, understanding the fail-
re mechanisms and the complex interplay of process and material
arameters in the failure occurrence in metal forming operations
as attracted the attention of many researchers for more than six
ecades. The knowledge of different failure mechanisms is crucial
o improve product quality and design methodologies.
Barrett (1997) considered in his report about the fasteners that
he cold heading (CH) process is one of the most important multi-
tage metal forming processes because of its many advantages over
achining of the same part, including high productivity for com-
lex final shapes, minimum material waste and increased tensile
trength from cold working. Chitkara and Bhutta (2001) who stud-
ed the near-net shape heading of splines and solid spur gear forms
∗ Corresponding author. Tel.: +1 514 398 8546; fax: +1 514 398 7365.
E-mail address: amar.sabih@mcgill.ca (A. Sabih).
924-0136/$ – see front matter © 2008 Elsevier B.V. All rights reserved.
oi:10.1016/j.jmatprotec.2008.11.023
properties and their microstructure.
© 2008 Elsevier B.V. All rights reserved.
reported that the decision to use the CH process to manufacture
certain products depends on the complexity in the shape of the
head and the material employed and the possibility of internal and
external failure occurrence.
Currently, the CH industry favors using faster headers, reducing
the number of manufacturing stages, and producing high-strength
fasteners without final heat treatment. To achieve these goals, mod-
ified process designs are required which result in higher strains and
strain rates, which cause two familiar types of ductile defects in the
cold-headed products. The first is the external oblique or longitu-
dinal crack caused by the exhaustion of the material ductility. The
second failure was reported by Bai and Dodd (1992) in their study
for the adiabatic shear band (ASB) phenomenon. They reported that
the ASBs may lead to internal crack. Okamoto et al. (1973) stated in
their study of different forming processes that this type of failure
may result in splitting of the fasteners’ heads as shown in Figs. 1–3,
respectively.
Many researchers have focused on surface defects. Cockroft and
Latham (1968) and Lee and Kuhn (1973) studied this defect in upset-
ting and presented different criteria to predict it. Nickoletopoulos
http://www.sciencedirect.com/science/journal/09240136
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mailto:amar.sabih@mcgill.ca
dx.doi.org/10.1016/j.jmatprotec.2008.11.023
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A. Sabih, J.A. Nemes / Journal of Materials
2000) studied the surface defect in CH process and reached to a
onclusion that Cockroft–Latham criterion can be used to predict
he surface defect in CH process with a reasonable accuracy. In
ddition to the large number of studies that covered this type of
ig. 1. Cold-headed industrial bolt: (a) without defect, using material 1, (b) shank of the d
material 2), and (d) magnified detail of the ductile crack of the separated industrial bolt h
ig. 2. Sectioned bolts revealing material flow inside the bolt head after etching with Fry’
efect bolt (material 1), (b) with defect bolt (material 2).
ssing Technology 209 (2009) 4292–4311 4293
failure extensively, it can be visually inspected and accounted for
in the process design of the component for removal by means of
machining or trimming. Thus far, internal ductile failure due to the
ASB phenomenon in CH process has not received the same atten-
efective bolt (material 2), (c) separated industrial bolt head due to an internal crack
ead (material 2).
s reagent (cupric chloride 36 g; 145 ml hydrochloric acid; 80 ml water): (a) without
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294 A. Sabih, J.A. Nemes / Journal of Materials
ion. The clear need for an in-depth understanding of the internal
uctile failure mechanisms which limits the development of new
H designs and the improvement of cold-headed product quality
otivated this work.
To have a full understanding of the complicated parameters and
echanisms contributing in the initiation and propagation of the
nternal ductile failure, the first part of this paper will provide a
etailed introduction covering important details about the CH pro-
ess and the ASB phenomenon followed by examples of ductile
ailure in cold-headed parts. The second part of this paper will focus
n the experimental and metallographic procedures used to inves-
igate the failure initiation and propagation. The last part of this
aper presents the detailed result and discussion.
.1. Cold heading process
The term heading refers to cold, warm, or hot working of metal
mploying a static or dynamic impact compressive force causing the
etal at one end of the rod or blank to form a head. Yoo et al. (1997)
efined the cold heading (CH) process as a forming process that is
erformed without an external heat source and it involves applying
force to the free end of a metal workpiece contained between a
ie and a punch, by one or several blows of the punch. The CH pro-ess is a high-rate deformation process with strain rates exceeding
00 s−1. Kawashima (1992) reported that the CH process is used
o produce a large variety of components such as fasteners, studs,
mall shafts, and other machine parts. The automotive, construc-
ion, aerospace, railway, mining and electrical product industries
re major consumers of such parts.
Matsunaga and Shiwaku (1980) concluded in their study of man-
facturing CH quality wire rods and wires that material properties
nd process parameters significantly affect the heading results.
oreover, the CH material must meet two sets of requirements.
he first set concerns the material’s ability to be formed with-
ut failure while the second set is linked to the properties of the
nal product. Sarruf et al. (1999) stated that in order to form
ithout failure, CH wire should be free from surface defects and
nternal inclusions. Nickoletopoulos et al. (2001) reported that the
hemical composition of CH wires plays a significant role in the
aterial ductility and the final properties of the products. Low- and
edium-carbon steels (0.06–0.30% C), are commonly used in the
H process, as they offer good ductility. Also, Nickoletopoulos et al.
2001) referred that the CH material microstructure is usually com-
rised of a ferrite matrix with varying amounts of lamellar pearlite.
ig. 3. Cold-headed brass bolt experienced ductile failure in service along the ASB: (a) an
ith permission of ASM International (www.asminternational.org).
ssing Technology 209 (2009) 4292–4311
A spheroidized microstructure provides the desired ductility for
CH of complex geometries. The required final strength properties
are usually achieved through alloy addition and/or heat treatment
(quenching and tempering).
1.2. ASB phenomenon
Wright (2002) defined the ASB as a two-dimensional, narrow,
nearly planar region of very large shearing that occurs in metals and
alloys experiencing intense dynamic loading (as in the CH process).
When the band is fully developed, the two sides of the region are
displaced relative to each other. However, the material still retains
full physical continuity from one side to the other. The thickness
of the most heavily sheared region might be few tens of microns
or less, and its length might extend many millimeters or centime-
ters. Generally, ASBs form under impact loading at high-strain rates
(higher that 102 s−1) and high strains. With increasing plastic defor-
mation, the work hardening (from increases in strain and strain
rate) results in an increase in the flow stress. However, most of
the plastic work (90–95%) is converted into heat causing a local
temperature increase and a flow stress decrease due to thermal or
work softening. Thus, a competing mechanism between the work
hardening and the thermal or work softening commences and con-
tinues in the deformation zone. As the work hardening mechanism
dominates over the work softening at the beginning, an increase in
the flow stress occurs and with continuing deformation, the ther-
mal or work softening mechanism can progressively dominate over
work hardening increases, which then triggers unstable deforma-
tion. Wright (2002) also reported that this instability condition
will force the deformation to localize into a narrower band that
through further localization can lead to final failure. Cowie et al.
(1989) showed in their study of microvoid formation during shear
deformation of ultrahigh-strength steels that this failure initiates
and develops by progressive nucleation, growth and coalescence
of microvoids and microcracks. A sharp drop in the load-carrying
capability for the deformed zone will associate with microvoid for-
mation causing what is known as the microvoid softening. Once the
failure starts to develop, the combined work softening effect of the
thermal and microvoid softening mechanisms will compete with
the work hardening effect within the band.
Zukas (1990) reported that there are two types of ASBs, namely
the deformed adiabatic shear bands (DASBs) and the transformed
adiabatic shear bands (TASBs). DASBs occur in materials that do
not undergo phase transformation, or when the local temperature
d (b) sectioned failed bolt, (c) crack surface in a failed bolt (Reitz, 2005) Reprinted
http://www.asminternational.org/
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A. Sabih, J.A. Nemes / Journal of Materials
ithin the band is not high enough to cause phase transformation.
he damage and fracture process in DASBs involves a number of
etallurgical events involving different steps between void nucle-
tion and crack propagation. The initial phase of damage coincides
ith void nucleation at inclusion edges or at grain boundaries due
o the interaction of local stresses and dislocations phenomena. The
amage in the material is then driven by the accumulated plastic
train and affected by the stress triaxiality. Hambli (2001) reached
n important conclusion that deformation under compressive (neg-
tive) stress triaxiality increases material formability in comparison
ith deformation under tensile loading state, i.e. positive stress tri-
xiality. Zukas (1990) declared that the competition between the
tress triaxiality levels reached in the specimen and the high-plastic
trains will determine the failure site.
Daridon et al. (2004) reported that the localization of plastic
eformation in narrow shear bands is a major damage mechanism
hat leads to ductile failure during high-strain rate deformation. Lee
t al. (2004) reported that internal cracks occur within the shear
Fig. 4. As-received testing materials microstructure: (a) 1008 steel, (b) 101
ssing Technology 209 (2009) 4292–4311 4295
band, and become the sites of eventual failure. Venugopal et al.
(1997) concluded that this can take place when the material is sub-
sequently subjected to dynamic impact loading during forming or
in service.
TASBs occur in materials that undergo phase transformations,
and are often found in high-strength steel in locations where the
critical temperature for the transformation of ferrite to austenite
(Ac3) is exceeded. After high-strain rate forming, the rapid heat dis-
sipation from the ASB to the surrounding matrix material results
in a fast decrease in temperature below the Ms limit. This sud-
den change in temperature leads to austenite transformation into
martensite, yielding the TASBs. TASBs may fracture in a brittle man-
ner in directions normal to the local tensile stress. If the matrix is
sufficiently ductile, brittle fracture will be limited to the band, but
sometimes the fracture may extend to the matrix. Rogers (1983)
stated that even if the TASB does not fracture, it remains a brit-
tle fracture path in the middle of the ductile matrix and might
lead to catastrophic failure. Moreover, Klepaczko et al. (1988) have
8 steel, (c) 1038 steel, (d) 1541 steel, (e) 8640 steel, and (f) DP steel.
4 Processing Technology 209 (2009) 4292–4311
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Table 1
Chemical analysis of the testing materials (wt%).
AISI steel grade C Mn P S Si Cu Ni Cr
1008 0.09 0.4 0.018 0.012 0.01 Nil 0.01 0.02
1018 0.16 0.69 0.006 0.006 0.22 0.04 0.08 0.08
1038 0.38 0.82 0.007 0.002 0.19 0.04 0.08 0.08
1541 0.37 1.41 0.011 0.018 0.22 0.16 0.08 0.09
8640 0.40 0.91 0.009 0.005 0.25 0.14 0.42 0.43
DP 0.088 1.69 0.009 0.011 0.62 0.04 0.07 0.07
AISI steel grade Mo Sn Al N B V Co
1008 0.002 0.002 0.038 0.0046 – – 0.004
1018 0.001 0.002 0.002 0.0026 0.0001 0.005 –
1038 0.001 0.003 0.003 0.0044 0.0001 0.005 –
Method A requires a survey of a polished surface square area
of 160 mm2 at 100× magnification parallel to the longitudinal axis
of the wire. The severity level of the worst fields and the series
type of four inclusion types (A, B, C, and D) should be reported for
every specimen depending on Tables 2 and 3. Table 2 presents the
296 A. Sabih, J.A. Nemes / Journal of Materials
ndicated that shear bands either with or withoutphase transfor-
ations, are considered as a site of fracture initiation.
.3. Examples of internal ductile failure in cold-headed products
The following paragraphs present cold-headed products failed
y internal ductile fracture along the ASBs. Fig. 1 shows in detail
ndustrial bolts were made of two different 1018 steels from two
ifferent CH steel providers (material 1 and material 2) using the
ame CH sequence. The industrial bolt made of material 1 (Fig. 1(a))
id not exhibit internal cracks while the second bolt made of mate-
ial 2 failed by a crack resulting from the head splitting when a
ow-tensile load was applied. Fig. 1(b)–(d) shows the details of the
rack inside the head of the bolt made of material 2. The met-
llographic study of the sectioned bolts made of the material 2
Fig. 2(b)) confirmed the nucleation, growth and coalescence of
icrovoids leading to microcracks.
Reitz (2005) investigated the failure of brass bolts after 12 years
n service. The microscopic investigation revealed that the bolts
ailed by macrocracks along the highly deformed band inside the
olt’s head (Fig. 3).
In general, the ductile crack occurred inside a highly localized
egion of the ASB. The occurrence of ductile failure in material 1
sing the same CH multistage design in the part in Fig. 1 raises
mportant questions about the ASB phenomenon and how it leads to
uch a catastrophic failure. The ductile failure that took place inside
olt heads made of material 1 presents clear evidence that the cur-
ent rules of thumb are neither effective nor adequate in fulfilling
he needs of the new design developments in the rapidly expand-
ng CH industry. They do not provide enough guidance to fulfill the
eeds of new design developments in the CH industry. Therefore,
he CH industry is in need of systematic and theoretical studies to
nderstand the failure mechanisms and the different parameters
ontributing to failure initiation and propagation in cold-headed
roducts.
These examples of failed cold-headed products reveal the appar-
nt lack of knowledge about the internal ductile failure mechanism
nd the role of the ASB phenomenon in the presence of these defects
n cold-headed products
. Experimental
.1. Material selection and characterization
The CH quality AISI steel grades chosen for this research are:
008, 1018, 1038, 1541, 8640 and dual phase (DP) steel which cover
ow- and medium-carbon content steels and also provide different
H quality steels ranging from best to poorest CH formability steels.
he 1018, 1038, 1541, and 8640 steels were received in rod form
rom Ivaco Rolling Mills (Ontario, Canada), while the 1008 steel was
rovided from the undeformed part of the bolts made of material
.
In addition, DP steel was also included in the test materials as it
ossesses unique properties which show great potential in the CH
ndustry. Pierson (1986) reported that the DP steel is intended to
e cold headed directly from the “green rod” and reaches the final
roperties of the cold-headed products without heat treatment.
Optical microscopy revealed a microstructure consisting of fer-
ite matrix and lamellar pearlite for all steel grades except the DP
teel which consisted predominately of ferrite grains and a marten-
itic second phase. The as-received microstructure of all the testing
aterials is shown in Fig. 4. Table 1 lists the chemical composition
f the selected material.
Quasi-static tensile tests were performed on specimens
achined to ASTM-E8 standard at an approximate strain rate
1541 0.021 0.010 0.004 0.0058 0.0000 0.002 –
8640 0.217 0.010 0.005 0.0063 0.0000 0.032 –
DP 0.458 0.005 0.003 0.0035 – 0.007 0.033
of 0.002 s−1, the force and the axial displacement values were
recorded. The engineering stress–engineering strain curves were
constructed from the load–elongation measurements. These curves
were used to calculate the true stress–true strain curves (Fig. 5) for
all testing materials.
2.2. Testing material cleanliness
Cleanliness is a measure of the inclusion content and inclusion
types that are deleterious to processing performance and final prod-
uct properties.
Dodd and Bai (1987) indicate that the inclusions refer to non-
metallic particles that are held mechanically in the material matrix.
Such particles include oxides, sulfides or silicates. The presence of
these inclusions strongly influences the ductile failure mechanism.
Thus control of cleanliness and second-phase particle shape can
influence material ductility.
The inclusion rating process for this study was performed using
automatic image analysis inclusion ratings in accordance with
microscopic method A (worst fields) of E45 of the ASTM standard.
Worst field is a class of rating in which the specimen is rated for
each type of inclusion by assigning the value of the highest severity
rating observed of the inclusion type anywhere on the specimen
surface.
Fig. 5. The tensile quasi-static true-stress vs. true strain curves for the testing mate-
rials.
A. Sabih, J.A. Nemes / Journal of Materials Processing Technology 209 (2009) 4292–4311 4297
Table 2
Total inclusion length or number for minimum severity level numbers (method A) (ASTM E45-97).
Severity level Total length in one field at100×, min. (mm) Number of inclusions in one field
Type A Type B Type C Type D
0.5 3.7 1.7 1.8 1
1 12.7 7.7 7.6 4
1.5 26.1 18.4 17.6 9
2 43.6 34.3 32.0 16
Table 3
Inclusion width and diameter parameters (method A) (ASTM E45-97).
Inclusion type Thin series (T) Heavy series (H)
Width min.
(�m)
Width max.
(�m)
Width min.
(�m)
Width max.
(�m)
T
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Table 4
Inclusion rating test results for each testing material using ASTM E45-97 method
A.a.
AISI grade A B C D
1008
steel
None None None D1
None None None D1
1018
steel
None None C1 D1
None None None D1
None None None D1
1038
steel
None None C1 D1
A1 None None D1
1541
steel
A1 None C2 D1
A1 None None D1
A1 None None D1
8640
steel
A1 None None D1
A1 None C2 D1H
None None None D1H
DP
steel
None None C1 D1
None None C2 D1
None None C1 D1
ype A 2 4 >4 12
ype B 2 9 >9 15
ype C 2 5 >5 12
ype D 2 8 >8 13
inimum severity level number for each inclusion type and Table 3
rovides the inclusion dimensions for the thin (T) and heavy (H)
eries according to method A.
This microscopic method places inclusions into several
omposition-related categories: sulfides, oxides, and silicates. The
ypical chemical types of inclusions in Type A are sulfide stringers,
hile Type C are silicate stringers. Type B are oxide stringers while
ype D is a globular oxide. Method A requires length measurement
f Type A inclusions, the total stringer length of Type B and C inclu-
ions, and the number of Type D inclusions. Type D globular oxide
ay not exceed an aspect ratio of 5:1.
The inclusion rating test results for the six testing materials are
isplayed in Table 4 and the morphologies of the inclusion types
re shown in Figs. 6–8.
A1 None None D1
a Cleanliness tests were performed by Ivaco Rolling Mills, L’Orignal, Ontario.
Fig. 6. Inclusion type A1 in the testing materials.
4298 A. Sabih, J.A. Nemes / Journal of Materials Processing Technology 209 (2009) 4292–4311
C1 in
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Fig. 7. Inclusion type
Zhang and Thomas (2002) reported that the inclusion size is
ery important because large inclusions are the most harmful to the
echanical properties of CH materials. Sometimes a catastrophic
efect is caused by a single large inclusion in an entire steel heat.
hus, producing clean steel requires not only controlling the mean
nclusion content in the steel but also avoiding inclusions larger
han the critical size.
.3. Tests
The drop weight compression test (DWCT), as described by
ickoletopoulos (2000), facilitates the testing of the CHP and
s capable of generating ASBs during upset testing. The DWCT
achine, as illustrated in Fig. 9, consists of a tower enabling inter-
hangeable weight plates to be dropped from heights up to 2.4 m.
he die set configuration (Fig. 10) rests on a centralcolumn that is
elded to the base of the DWCT machine. Specimens for CH were
achined with a tolerance of 0.02 mm from as-rolled rod material
o a cylindrical configuration of 5.3 mm in diameter with aspect
atios of 1.6 and 1.8. A series of tests is performed by varying the
the testing materials.
weights until internal fracture is found at different locations inside
the ASB.
A guided cylindrical-pocket die set was designed for the DWCT
machine. A sleeve guides the dies and reduces die movement during
testing. Air vents in the sleeve prevent pressure build-up. An impact
plate is used on top of the upper die to prevent direct impact with
the die. Fig. 10 presents a schematic assembly of a test specimen
which is placed between the upper and lower dies. This assembly
rests atop the force sensor inside a cylindrical pocket.
2.4. Sample preparation
The cold-headed samples were sectioned transverse to the head-
ing direction using a high-concentration diamond wafering blade
on a slow speed (200 rpm) cut off saw that was operated with an
oil coolant. Sectioned specimens were mounted in Bakelite and
prepared for metallographic examination using automated tech-
niques for grinding and polishing. Specifically, the specimens were
ground with successive papers of SiC from 280 grit to 600, fol-
lowed by rough polishing using a 9 �m diamond suspension with
A. Sabih, J.A. Nemes / Journal of Materials Processing Technology 209 (2009) 4292–4311 4299
D1 in
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Fig. 8. Inclusion type
n alcohol based lubricant on a silk cloth and final polishing with 3
nd 1 �m on a nylon and porous pad, respectively. After ultrasonic
leaning, the specimens were etched by immersion in a 2% nital
olution.
Microstructural examination was performed using optical light
icroscopy with polarizing capabilities and image analysis soft-
are to examine the overall characteristics of the ASBs and the
oids and cracks initiated in the interior or along the ASBs. High-
esolution imaging for examining the topography of the voids
nd cracks in the ASBs was performed using scanning elec-
ron microscopy operated in secondary electron (SE) mode and
quipped with an X-ray detector (EDAX) for elemental analysis of
econdary phases. Mapping the topography of the voids and cracks
n the ASBs was performed using an Olympus OLS1200 laser scan-
ing confocal microscope (LSCM) equipped with an argon laser
488 nm) and objective specifications of up to 100× with a 6× opti-
al zoom. As the image produced by scanning the x- and y-axes
ith a spot laser beam in the LSCM originates from a shallow focus
epth, for the DP steel specimens a series of precisely focused opti-
the testing materials.
cal images were acquired along the z-axis of the microscope. The
successive images were then overlapped using specialized recon-
struction software to obtain a three-dimensional extended focus
image of the cross-sectional view that contained both height and
intensity information. Specifically, the topographic features in the
extended focus image of the ASB structure enabled the characteri-
zation of the size of the width and depth of the voids and cracks.
3. Results and discussion
In general, microscopic examination and FE analysis of the DWCT
specimens at various regions of each specimen during different
deformation levels showed that the deformation concentrates in
the ASBs (Sabih et al., 2005, 2006a,b). The microstructure of the
DASBs is laminar consisting of highly elongated grain layers of fer-
rite and pearlite or second-phase particles (Fig. 11). Voids and cracks
found along the ASBs inside DWCT specimens (Fig. 12) coincided
with the cracking that occurs inside the head of the industrial bolt
shown in Fig. 1. This finding agrees with that of Qiang and Bassim
4300 A. Sabih, J.A. Nemes / Journal of Materials Proce
Fig. 9. Schematic of the DWCT machine.
Fig. 10. DWCT specimen inside the die set arrangement.
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ig. 11. The microstructure of the DASB (1541 steel, aspect ratio of 1.6, drop height
f 2.4 m and drop weight 29.5 kg).
2003) who studied the effect of process parameters on the forma-
ion of the ASBs in metals and observed the initiation and growth of
racks along the ASBs. Similar observation was reported in the work
f Nemat-Nasser et al. (2001) who reported in their studies of ASB
ig. 12. Microscopic views showing the internal cracks and/or voids along the ASB
DP steel, aspect ratio = 1.8, dropped weight = 33.8 kg, height = 2.4 m).
ssing Technology 209 (2009) 4292–4311
phenomenon and the associated failure that these bands usually
precede the failure mechanism for most metals and alloys.
Ductile fracture occurs progressively by void nucleation and
growth mechanisms with increasing plastic strain. Different mech-
anisms of ductile failure have been reviewed in many works: Dodd
and Bai (1987), Dodd and Atkins (1983) and Garrison and Moody
(1987). All these reviews consider the most important factor affect-
ing the material formability is the number and morphology of
inclusions and second-phase particles.
Dodd and Atkins (1983) found in their study of flow localization
in shear deformation of void containing and void-free solids that
ductile failure occurs following this sequence of events:
(1) Nucleation of voids occurs by decohesion at second-phase parti-
cles or at inclusion–matrix interfaces, or by particle or inclusion
cracking. Dodd and Hartley (1991) concluded that in certain
cases, these voids can nucleate at grain boundaries or grain
boundary triple points.
(2) An increase in the volume fraction of voids with increasing
applied plastic strain.
(3) At a critical strain, plastic deformation is localized along shear
bands between the voids.
(4) Cracks occur along the bands of intense shear, forming what
Rogers (1959) calls “void sheets” mechanism of crack propaga-
tion.
Masheshwari et al. (1978) studied quality requirements for steel
CH grades and concluded that the second-phase particles and inclu-
sions, such as heavy silicate, sulphide, alumina and globular oxide
inclusions, should not exceed the acceptable limits of the cold head-
ing material. Thomson (1990) pointed out that these particles and
inclusions act as stress raisers and can support microvoid nucle-
ation, growth, and coalescence, leading to ductile fracture.
Meyers et al. (2001) reported that the localization of the ASBs
is an important deformation mode that causes ductile fracture in
metals. Wei et al. (2006) concluded in their study of mechanical
behavior and dynamic failure under uniaxial compression that the
thermal and geometric softening and material flow patterns are
responsible for dynamic failure under compressive loads.
Ductile failure caused by ASBs within the DWCT specimens
under compressive loads is a complicated process which depends
on the interplay of many factors. The following sections will detail
the ductile failure mechanisms, the void nucleation, coalescence
and crack initiation mechanism. Furthermore, the interplay of
inclusion types and content, material microstructure, material flow
pattern and the localization phenomenon will be discussed along
with the impact of these factors on the occurrence of ductile failure.
3.1. Void initiation mechanism
A careful reading of the literature shows that there is a debate
about the issue of void nucleation under a compressive state of
stress. Petersen et al. (1997) reported that ductile failure is unlikely
to occur under a compressive stress state. Teng et al. (2007) stud-
ied the transition from adiabatic shear banding to fracture and
reported that ductile fracture under high-negative stress triaxial-
ity cannot occur. On the other hand, many researchers have found
that inclusions or inter-metallic particles serve as void nucleation
sites under compressive loads: Puttick (1959) showed in his study
of ductile fracture that microvoid initiation and growth is caused
by the debonding of second-phase particles produced by deforma-
tion in the surrounding matrix.Cox and Low (1974) found that void
nucleation and growth occurs more readily at the larger inclusions.
Rogers (1960) explained the role of the small and large particles
and inclusion in the initiation and growth of voids leading to a duc-
tile fracture. Similarly, Gurland (1972) also explained the effect of
A. Sabih, J.A. Nemes / Journal of Materials Processing Technology 209 (2009) 4292–4311 4301
F Z. Sol
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ig. 13. DWCT specimen of 1018 steel showing the locations of the CASB, UDZ, and LD
epresent the shear direction and the small arrow shows the material flow pattern.
rop weight 22.5 kg).
he inclusions on the nucleation, growth, and coalescence of micro-
copic voids causing fracture of ductile solids. Horstemeyer and
okhale (1999) modeled void–crack nucleation process for duc-
ile metals with second-phase particles. They concluded from the
odeled compression, tension, and torsion experiments that the
oid–crack nucleation can take place under compressive loading
ondition.
Nemat-Nasser et al. (2001) also reported that cracks of this kind,
hich are produced by local defects, can be created by a local tensile
tress state under an overall compressive applied stress. They also
ound that the occurrence of tensile stresses in the vicinity of the
nclusions and/or second-phase particles create the classical con-
itions for void nucleation and growth. Thus, fracture initiation in
SBs under compressive loading of the DWCTs is highly probable.
The present findings are supported by previous work by Dodd
nd Atkins (1983) showing that void growth is sensitive to a
ydrostatic stress state. Furthermore, they found that voids nucle-
te at second-phase particles or inclusions within the ASB under
ompressive hydrostatic stresses. Under compressive loading con-
itions with a local tensile stress state, both void nucleation and
rowth can occur in steels, albeit at a lower rate in comparison to
he dominating failure mechanisms of void formation and propaga-
ion under tensile loading. Park and Thompson (1988) have shown
hat void nucleation occurs at both sides of the tensile pole of a
article under compressive loads. Horstemeyer and Gokhale (1999)
ig. 14. Material flow contours for various regions on a DWCT specimen (drop weight =
ocated at the center of the DWCT specimen (b) the region adjacent to the UDZ, (c) the r
WCT dies (25×), (e) the region deformed within the DWCT dies (100×), (f) the region in
id lines represent imaginary boundaries of the ASB and the UDZ and LDZ. Big arrows
ashed arrow represents the mid of the CASB (aspect ratio = 1.8, drop height = 2.4 m,
reported the possible occurrence of void nucleation even under
compression loads as tensile states exist for compression loadings.
3.2. Effect of material flow pattern on void nucleation
Void nucleation around second-phase particles and non-
metallic inclusions is highly influenced by the nature of the material
flow during deformation. Sarruf et al. (1999) reported that the metal
flow pattern during the CH process supports the void nucleation
mechanism at non-metallic inclusions and the surrounding mate-
rial. Meyers et al. (2001) reported that the material flow pattern
during shear localization is an important factor in sub-grain break-
age and rotation. Tvergaard (1981) reported that this mechanism is
responsible for void nucleation due to brittle cracking or decoher-
ence of inclusions.
Detailed microscopic examination of various regions of the
DWCT specimens showed that the deformation is concentrated in
the localized central ASBs (CASB) with well-defined flow contours
starting from the upper dead zone (UDZ) and the lower dead zone
(LDZ), as illustrated in Fig. 13. Specifically, there is a sharp transi-
tion in the behavior of the material flow as observed by the contour
lines which traverse from the dead zones to the CASB (Fig. 14(a)), to
the CASB edge at the mid-axis (Fig. 14(b) and (c)). In the deformed
region within the DWCT dies, the material flow behavior showed
elongated semi-elliptical contours that increasingly concentrated
35.5 kg and height = 2.4 m) showing: (a) the dead zone region adjacent to the ASB
egion near the specimen edge at the mid-axis, (d) the region deformed within the
the CASB showing the merging of two dead zones and (g) the center of the CASB.
4302 A. Sabih, J.A. Nemes / Journal of Materials Proce
F
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ig. 15. Excessive deformation at the region between the UDZ, the LDZ and the
ASB resulting in microscopic crack formation at the boundary between the top and
ottom dead metal zones and at the inflection point of material flow (DP steel, aspect
atio of 1.6, drop height of 2.4 m and drop weight 31.6 kg).
owards the mid-axis of the specimen, as shown in Fig. 14(d) and (e).
t the center of the DWCT specimen, the CASB zone was observed
o have highly deformed material with voids (Fig. 14(f) and (g)).
ith further deformation, the region extended along (f), (g), (b)
nd (c), and was observed to have excessive deformation result-
ng in microscopic crack formation at the boundary between the
UDZ) and the (LDZ) and at the inflection point of material flow
Fig. 15).
An analysis of the void characteristics, using the LSCM, indi-
ated that an increase in the plastic strain caused both an increase
n the void density and size, as illustrated in Fig. 16. Thus, in
he CASB region, longer cracks were observed with increasing
rop weight for the DWCT specimens. The occurrence of the long
racks in the CASB was usually observed in two zones. The first
one was along the center of the CASB (Fig. 14(f)), and the other
ig. 16. A comparison of void density inside the ASB of two specimens (DP steel,
spect ratio of 1.8 and drop height of 2.4 m) that were deformed with a drop weight of
a) 38 kg and (b) 31.5 kg. The dark heights on the surface represent the void locations
nd their inverted depth.
ssing Technology 209 (2009) 4292–4311
was found in the upper and lower regions adjacent to the CASB
(Fig. 14(a)–(c)).
3.3. Void nucleation mechanism at the inclusions–matrix
interface
Siruguet and Leblond (2004a,b) concluded that the void nucle-
ation, growth and coalescence mechanism is highly dependant on
the stress state under which deformation takes place. Fig. 17 dis-
plays possible void nucleation mechanisms under different stress
states in order to have a better understanding of the ductile fail-
ure occurring inside the ASB. When the stress triaxiality is high
(tensile loads/positive), voids grow in the directions of the tensile
loads (Fig. 17(b) and (c)), and the enclosed inclusions do not influ-
ence their growth. Moreover, it has been reported by Pardoen and
Delannay (1998) that the initially spherical voids will change to
ellipsoidal voids under tensile triaxiality.
Siruguet and Leblond (2004a,b) showed that under compres-
sive loads from all directions, or negative stress triaxiality, voids
can be subjected to compression in one or several directions,
thereby changing their shape. Also, if they are still in contact
with the enclosed inclusions in these directions, inclusions pre-
vent their shrinkage, a process known as void locking illustrated in
Fig. 17(d). Pardoen and Delannay (1998) reported that under low-
compressive stress triaxiality, the ductile matrix may flow around
the inclusion resulting in the failure of the interface forming small
voids at the sides of the inclusion (Fig. 17(e)). Conversely, under
high-compressive loads, the fracture of the inclusion is possible
(Fig. 17(f)).
In the case of shear banding under compressive loads, the forma-
tion of microvoids has been reported by Pirondi and Bonora (2003)
and Sabih et al. (2005, 2006a). They indicated that the microvoids
have smaller dimensions than the tensile cases. They also con-
cluded that voids under compressive loads do not increase their
dimensions but stretch and eventually coalesce by internal necking,
known as microvoid sheeting (Fig. 17(g)).Evidence of void and crack initiation and propagation at the non-
metallic inclusions and stringers sites was put forward in the work
of Vora and Polonis (1976). They found that the weakly bonded
particles to the matrix had elongated in the direction of material
flow. Once decohesion between the matrix and the inclusion or
second-phase particle occurred, the nucleated voids were observed
to grow with increasing strain from an initial spherical shape to an
ellipsoidal form or sheet form, as illustrated in Fig. 17(h). Boyer et
al. (2002) modeled void growth under different stress states and
concluded that void volume increases under compressive loads if
these voids are filled with inclusions.
Meyers et al. (2001) noticed that the different rates of material
flow around the inclusion supports decohesion and void nucleation
mechanisms between the matrix and the inclusion or the second-
phase particle as shown in Fig. 17(i).
In DWC testing, the deformation takes place under compressive
stress triaxiality over almost the entire specimen. The ASB occurs
inside the DWCT as a result of the softening mechanisms and the
shear deformation between the edges and the center of the band.
The combined compressive and shearing deformation results in
ductile failure inside the band. The void nucleation and coalescence
follows some of the illustrated mechanisms in Figs. 17 and 18.
Fig. 18 displays some metallographic observations of the void
nucleation mechanisms around the spherical inclusions (Type D).
Fig. 20(b)–(f) show the steps of the void sheeting process inside the
CASB of the 1541 steel. This process starts when a void nucleates at
the edges of the inclusion due to the material flow around it. With
more deformation, the inclusion changes its shape from spheri-
cal to ellipsoidal and the void propagates along the material flow
direction forming a void sheet (Fig. 18(e)). This void sheet occurs
A. Sabih, J.A. Nemes / Journal of Materials Processing Technology 209 (2009) 4292–4311 4303
Fig. 17. Illustrations of the void and crack nucleation mechanisms around the inclusions under different stress states: (a) inclusion inside the undeformed material (no void);
(b) under multiaxial tensile loads, void growth in all directions; (c) under uniaxial tensile load, void nucleation in the load direction; (d) under multiaxial compressive loads,
void shrinks on the inclusion (void locking); (e) under uniaxial compressive load, void nucleation at the sides of the inclusion; (f) breakage of inclusion under multiaxial
c on un
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ompressive loads; (g) void sheet nucleation and growth at the sides of the inclusi
he material flow direction under combined compression and shear stresses; (i) de
ifferent rates of material flow around the inclusion.
n a manner similar to the void sheeting mechanism illustrated in
ig. 17(h).
The metallographic inspection showed that some of the small
oids change their shape to an ellipsoidal geometry, while the large
nclusions were not affected by material deformation and contin-
ed to remain in their spherical shape. However, it was noticed that
oid sheets initiated at the vicinity of the big inclusions and propa-
ated along the material flow direction (Fig. 18(g)). These findings
re supported by the results of the ductile damage finite element
odel used by Boyer et al. (2002) to predict void growth under
ormal mean stress.
Due to the different rates of material flow around the inclusions,
he big spherical inclusion rotates in the direction of the high-flow
ate. This results in decohesion at the interface between the inclu-
ion and the surrounding matrix. These findings are supported by
imilar observations made by Meyers et al. (2001), who reported
hat the material flow pattern during shear localization is an impor-
ant factor in the breakage and rotation of inclusions.
Sabih et al. (2005, 2006a) showed in their metallographic inves-
igations and FE simulations that at low-plastic strains (less than 1)
nd low-negative stress triaxiality, voids were observed to initiate
round the elongated inclusions (Fig. 19(b)). Bandstra et al. (2004)
der compressive and shear loads; (h) fractured or elongated weak inclusion along
ion between the inclusion and the matrix due to rotation of inclusions caused by
reported void nucleation at the non-metallic inclusion–matrix
interface at small strains. In this work, inclusions were also
observed to fracture during deformation (Fig. 19(c)). According to
the microscopical examination of fracture surfaces made by Broek
(1973), this can cause stress concentrations promoting ductile frac-
ture.
Type A and C (long inclusion) stringers are the favored places for
void nucleation and growth within the ASBs. The metallographic
study of the DWCT specimens showed that the void sheeting pro-
cess took place by decohesion at the stringer–matrix interface in
several steps. The void sheeting mechanism around the stringers
is the result of the material flow and strain localization within the
ASB. As shown in Fig. 14, the stringers move along the material flow
direction and change their alignment by approximately 90◦. This
sharp change in the flow direction at the early stages of deformation
triggers the decohesion process at the stringer–matrix interface at
the boundaries of the CASB (Fig. 14(b) and (c)).
When these stringers join the localized CASB, the softening
mechanisms and the high-strain rate deformation within the band
support the void sheeting process along the stringers sides and
around the spherical inclusions (Fig. 14(f) and (g)). Fig. 20(a) dis-
plays the sharp change in the stringers’ direction which supports
4304 A. Sabih, J.A. Nemes / Journal of Materials Processing Technology 209 (2009) 4292–4311
Fig. 18. Changes in the inclusion morphology and void sheet nucleation mechanism around inclusions at different deformation levels in 1541 steel: (a) spherical inclusion
inside the undeformed material, (b) inclusion shape changed from a spherical to an ellipsoidal form inside the DASB, (c) elongated inclusion inside the DASB, (d) void sheet
initiated at the edges of the elongated inclusion, (e) void sheet growth inside the DASB, (f) a large number of void sheets inside the highly deformed TASB, (g) void sheet
propagating ahead of the inclusion, and (h) rotation of the inclusion.
A. Sabih, J.A. Nemes / Journal of Materials Processing Technology 209 (2009) 4292–4311 4305
Fig. 19. The morphology of inclusions: (a) spherical inclusion in as-received (undeformed) DP steel, (b) elongated particles with adjacent voids initiated at low-plastic strains,
and (c) fractured elliptical inclusion.
Fig. 20. Void sheet mechanism around the stringers: (a and b) 1038 steel, aspect ratio of 1.8, drop height of 2.4 m and drop weight 26.2 kg, (c) 1018 steel, aspect ratio of 1.6,
drop height of 2.4 m and drop weight 28.5 kg, (d) 8640 steel, aspect ratio of 1.6, drop height of 2.4 m and drop weight 29.5 kg, and (e) 1541 steel, aspect ratio of 1.6, drop height
of 2.4 m and drop weight 23.4 kg.
4 Proce
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306 A. Sabih, J.A. Nemes / Journal of Materials
he void sheets inside the DWCT of 1541 steel. Usually, the voids
ucleate around the stringer, forming the void sheet inside the
eformed region. With further deformation, the void sheet, which
s considered as another softening mechanism, propagates along
he localized CASB (Fig. 20(a) and (b)).
In addition to void sheeting around the stringers within the
ASB, metallographic inspection revealed that these stringers are
he source of ductile failure inside the region deformed within the
WCT dies (Fig. 20(a) and (c)). In this region, the material flow pat-
ern forces the stringers to take the shape of semi-elliptical contours
hat become increasingly concentrated towards the mid-axis of the
pecimen: Fig. 20(a), (d) and (e). The stringers and the surrounding
aterial at these locations deform under low-compressive stress
riaxiality in comparisonwith the center of the CASB. Under this
tress state, and with further deformation, cracks initiate at the tip
ig. 21. SEM micrographs display the void nucleation at the ferrite–cementite interface
earlite (ferrite–cementite) in the undeformed zone (b) elongated pearlite in the directi
ucleation at the grain boundaries, (e) void nucleation at the ferrite–cementite interface
1038 steel, aspect ratio of 1.8, drop height of 2.4 m and drop weight 33 kg).
ssing Technology 209 (2009) 4292–4311
of the semi-elliptical stringers due to the stretch-like deformation.
This stretching occurs because the flow of the semi-elliptical sides
is restricted, due to the friction with the die faces, while the mate-
rial continues to push the tip of the semi-elliptical stringers away
from the specimen center.
3.4. Void nucleation mechanism at the grain boundaries and the
second phase–matrix interface
As presented in Section 2.1, the materials tested are
ferritic–pearlitic steels, except for the DP steel which has a
ferrite–martensite microstructure. The pearlite grains consist of
cementite lamella in a ferrite matrix. This type of microstructure
makes the materials subject to a new source of void nucleation
during plastic deformation under compressive loading.
, at the grain boundaries and around the inclusions inside the DASB: (a) lamellar
on of material flow, (c) void nucleation at the ferrite–cementite interface, (d) void
and at the grain boundaries, and (f) void nucleation around an elongated inclusion
A. Sabih, J.A. Nemes / Journal of Materials Processing Technology 209 (2009) 4292–4311 4307
Table 5
Inclusion rating result with the minimum inclusion length or number according to the minimum severity level numbers (method A) (ASTM E45-97).
AISI steel grade A1 Min. length (mm) C1 Min. length (mm) C2 Min. length (mm) D1 Min. number
1008 None None None D1 4
1018 None C1 7.6 None D1 4
1038 A1 12.7 C1 7.6 None D1 4
1
8
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D nd 14
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541 A1 12.7 C1 7.6
640 A1 12.7 C1 7.6
P A1 12.7 C1 7.6
1H is a heavy series of Type D1 inclusion, the inclusion width range is between 8 a
Cox and Low (1974) referred in their study of ductile fracture of
ISI 4340 steel that this failure is caused by void sheets composed
f small voids nucleated at the matrix–cementite (second-phase
articles) interface. Gurland (1972) studied the formation of cracks
nd voids at second-phase particles (carbides) under tension, com-
ression and torsion. Hahn and Rosenfield (1975) reported that void
rowth and coalescence was observed the cracked inclusions and
he grain boundaries. Nieh and Nix (1980) explained void growth
nd coalescence at grain boundaries. In conclusion, non-metallic
nclusions, second-phase particles, and grain boundaries are con-
idered to be the best candidate sites for void nucleation and ductile
ailure.
Tanguy et al. (2006) have also reported that the presence of
econd-phase particles, such as lamellar cementite or martensite,
re a source of void nucleation. Chen et al. (2002) showed that
amellar cementite has a delayed deformation rate as compared to
errite. Valiente et al. (2005) explained that the lamellar cementite
f a pearlitic grain has a much higher stiffness than ferrite.
Valiente et al. (2005) showed that due to microstructural
spects, the difference in the local flow behavior can lead to cemen-
ite elongation and fragmentation. This triggers void nucleation at
he discontinuities produced by the breakage of cementite and by
ecohesion of atomic bonds at the interface between the matrix
nd inclusion or second-phase particles. The voids grow, assisted
y plastic deformation, until they link up. Fragmentation of cemen-
ite within ASBs has also been reported by Chen et al. (2003). Argon
1976) found that voids were more easily nucleated at the larger
ementite particles than at the smaller ones.
Beyond the early stages of deformation in DWCT specimens,
etallographic study by SEM microscopy showed that the voids
ucleated at the cementite–ferrite interface, the pearlite–matrix
nterface and at grain boundaries (Fig. 21).
ig. 22. Ductile failure in DWCT specimens of 1008 steel: (a) a long crack along the CASB (
atio of 1.6, and drop weight 26.2 kg).
C2 32 D1 4
C2 32 D1, D1H 4
C2 32 D1 4
�m. The minimum width of the other inclusion types is 2 �m.
Within the matrix, the cementite inside the pearlite grains far
off the shear band have a lamellar shape as shown in Fig. 21(a).
Lamellar cementite inside the pearlite on the shear band side is
elongated along the material flow direction (Fig. 21(b)). At the
high-deformation zone close to the CASB, broken lamellar cemen-
tite was noticed, and clear void nucleation took place at the
ferrite–cementite interface (Fig. 21(c)).
Voids and microcracks nucleated at ferrite–pearlite interfaces
close to the localized plastic deformation zone of the CASB
(Fig. 21(d)). Closer to the highly localized plastic deformation within
the CASB, more void nucleation and cracking along the grain bound-
aries along the ASB was observed (Fig. 21(e)). These voids and
microcracks, along with the voids around the inclusions (Fig. 21(f))
contribute to ductile cracking along the ASBs.
3.5. Ductile failure characterization in the testing materials
The metallographic study of the DWCT specimens showed that
all materials experienced ductile failure by void nucleation and
coalescence, forming cracks along the CASBs. The ductile failure
of each material was the result of the contribution of all the
mechanisms of void nucleation at the inclusion–matrix interface,
second phase–matrix interface and at the grain boundaries, as
previously discussed. However, the level of contribution of each
mechanism in the final ductile failure varied depending on mate-
rial.
Ductile failure along the ASB inside the DWCT occurred under
compressive loads. The voids and cracks found in the failed zones
have the narrow, stretched, elongated shape or the void sheet shape
along the highly localized regions of the ASB.
In general, it is possible to classify the ductile failure into two
groups. The first group consists of materials that experienced very
aspect ratio of 1.6, and drop weight 37.6 kg), (b) void nucleation in the DASB (aspect
4 Proce
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308 A. Sabih, J.A. Nemes / Journal of Materials
arge cracks, including 1008 steel and 1018 steel. The second group
onsists of the material that experienced short ductile cracking
long the ASB. This group includes 1038 steel, 1541 steel, 8640 steel
nd DP steel.
The first group is composed of soft steels with low-inclusion
ontent. Table 5 shows that the only inclusion type in the unde-
ormed 1008 steel is the spherical inclusions (Type D1), while the
ndeformed 1018 steel contains the long stringer or inclusion Type
1 in addition to Type D1 inclusions.
Both steels experienced very long cracks which extended for
ore than 3 mm along the mid-axis of the CASB, as shown in
igs. 22(a) and 23(a) for 1008 steel and 1018 steel, respectively. The
EM metallographic inspection of the cracked specimens revealed
hat void nucleation and coalescence took place at the boundaries
f the highly elongated grain within the CASB in contribution to the
icrocracks at the ferrite–cementite interface. The metallographic
nvestigation of the role of the inclusions type in the 1008 steel
inclusion Type D1) showed that voids nucleated along the inclu-
ions leading to shorter cracks in comparison to the crack initiated
t the grain boundaries (Fig. 22(b)).
Similar conclusions can be made with respect to the 1018 steel
WCT specimens. The metallographic investigation showed that all
he void nucleation mechanisms contributed to the ductile cracks
ound within the highly localized deformed CASB in the DWCT
pecimens made of 1018 steel (Fig. 23). Void nucleation took place
ig. 23. Ductile failure in the DWCT specimens of 1018 steel: (a) a long crack along the CA
aspect ratio of 1.8, and drop weight 22.5 kg),and (c) void nucleation around the inclusion
ssing Technology 209 (2009) 4292–4311
around the inclusions (Fig. 23(b)), and at the interface between
both types of inclusions (Type C1 and Type D1) and the matrix.
The microscopic observation revealed that at the early stages of
deformation, void sheets and short cracks extend along the grain
boundaries in addition to microcracks at the ferrite–cementite
interface. With further deformation, the coalescence of these void
sheets and cracks form bigger cracks that extend over 3 mm along
the center of the CASB (Fig. 23(a)). The void nucleation along the
Type C1 inclusion is readily apparent in Fig. 20(c).
The materials in the second group can be characterized by the
occurrence of the TASBs that are accompanied with short cracks.
These cracks were noticed to be within, or at, the sides of the TASB
zone and the DASB zone of the band. As in the first group, all of
the void nucleation mechanisms contributed to the ductile cracks
found within the highly localized deformed and the transformed
part of the CASB. However, the long inclusions (Type A1, C1 and C2)
played an important role in ductile failure initiation as previously
discussed. The presence of inclusion type C2 in 1541 steel, 8640
steel and DP steel means that stringers with a total minimum length
of 32 mm are present in the matrix. As discussed previously, this
long inclusion is the favored location of void nucleation by the deco-
hesion mechanism at the inclusion–matrix interface. With further
straining inside the localized ASB, these voids coalesce around the
stretched long inclusion inside the band. Figs. 20(a), 20(b) and 24(a)
display the long cracks which initiated within the CASB of the 1038
SB (aspect ratio of 1.6, and drop weight 37.6 kg), (b) cracks at the grain boundaries
(aspect ratio of 1.6, and drop weight 20.2 kg).
A. Sabih, J.A. Nemes / Journal of Materials Processing Technology 209 (2009) 4292–4311 4309
Fig. 24. Crack initiation along the TASB: (a) 1038 steel, aspect ratio of 1.8, drop height of 2.4 m and drop weight 33 kg, (b) 8640 steel, aspect ratio of 1.8, drop height of 2.4 m
and drop weight 29.5 kg.
F eel: (
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ig. 25. Void coalescence in the ASB region of deformed DWCT specimen of DP st
5.6 kg).
teel, and shows the severe change of the material flow pattern.
his is known to cause the breakage of the long inclusion, provid-
ng another means of void sheet nucleation. Fig. 24(b) shows that
oids nucleated at the interface of the long inclusion–matrix in the
640 steel. Fig. 25 shows ductile failure along the ASB inside the
WCT specimen of DP steel. The ductile cracking mechanisms in
he 1541 steel are shown in Fig. 18.
Careful investigation of the microstructure at the side of the
ASB of the second group materials reveals that the elongated fer-
ite and pearlite grains also helped in the nucleation of narrow and
tretched void sheets along the localized band. Nevertheless, the
ole of the long inclusion clearly plays a much larger part in ductile
ailure.
a) 500× and (b) 1000× (aspect ratio of 1.8, drop height of 2.4 m and drop weight
4. Conclusion
The occurrence of internal ductile failure in cold-headed prod-
ucts presents a major obstacle in the fast expanding CH industry.
The current design rules of thumb failed in avoiding this type of
failure which results in reducing the material’s capability to carry
loads. This internal failure may lead to catastrophic brittle fracture
under tensile loads despite the ductile nature of the material.
The lack of specialized studies regarding the mechanisms of
internal ductile failure in cold-headed products and the failure of
the current CH design procedures to fulfill the needs of new devel-
opments in the expanding CH industry motivated this research.
A comprehensive testing and investigation methodology was
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310 A. Sabih, J.A. Nemes / Journal of Materials
sed to reveal the complicated interplay of process and material
arameters contributing to the initiation and propagation of the
nternal ductile failure in six CH quality AISI steel grades.
The metallurgical and microscopic studies showed that inter-
al ductile failure occurs progressively by void nucleation and
rowth mechanisms with increasing plastic strain inside the highly
ocalized CASBs. The most important factor affecting material
ormability is the number and morphology of any inclusions and
econd-phase particles. Moreover, the thermal and geometric soft-
ning mechanisms and material flow pattern are responsible for
ynamic failure under compressive loads.
It was confirmed that ductile failure occurs by a void nucleation
echanism which occurs by decohesion at second-phase particles
nd/or at inclusion–matrix interfaces, by particle or inclusion crack-
ng, or at grain boundaries. The localization process inside the CASBs
upported the void growth and coalescence along the band form-
ng narrow void sheets. Under compressive loading conditions, the
oid nucleation, growth and coalescence mechanism occurred due
o the local tensile stress state in the vicinity of the inclusions and/or
econd-phase particles.
The metallurgical investigations showed that the nature of the
etal flow pattern during the CH process, and the different rates of
aterial flow around the inclusions, support decohesion and void
ucleation mechanisms between the matrix and the inclusion or
he second-phase particle in the highly localize CASBs.
The void sheeting process starts when a void nucleates at the
dge of the inclusion due to the material flow around it. With more
eformation, the inclusion changes its shape from spherical to ellip-
oidal and the void propagates along the material flow direction
orming a void sheet.
Under combined compressive and shearing deformation, the
arge inclusions continued to retain their spherical shape. However,
t was noticed that void sheets initiated in the vicinity of the large
nclusions and propagated along the material flow direction. Type A
nd C long inclusions are the favored places for void nucleation and
rowth within the CASBs. The void sheeting process took place by
ecohesion at the stringer–matrix interface in several steps. The
harp change in the flow direction at the early stages of defor-
ation triggered the decohesion process at the stringer–matrix
nterface. The thermal softening and the high-strain rate deforma-
ion within the localizing band support the void sheeting process
long the stringers’ sides and around the spherical inclusions. With
urther deformation, the void sheet propagates along the localized
ASB.
In addition to non-metallic inclusions, the second-phase parti-
les and grain boundaries are considered to be the best candidate
ites for void nucleation and ductile failure. The difference in the
ocal flow behavior, due to microstructural aspects, can lead to
ementite elongation and fragmentation. This triggers void nucle-
tion at the discontinuities produced by the breakage of cementite
nd by decohesion at the interface between the matrix and inclu-
ion or second-phase particles.
The microstructure of the DASBs is laminar, consisting of narrow,
tretched, and highly elongated grain layers of ferrite and pearlite
r second-phase particles. Void sheets and cracks are found in the
ailed zones along the ASBs.
All materials tested experienced ductile failure by void nucle-
tion and coalescence, forming cracks along the ASBs. The ductile
ailure of each testing material was the result of the contribution
f all the mechanisms of void nucleation at the inclusion–matrix
nterface, second phase–matrix interface and at the grain bound-
ries. However, the level of contribution of each mechanism in the
nal ductile failure varied depending on the material.
In summary, a comprehensive experimental and metallurgical
tudy of the ASB phenomenon and the associated ductile failure
ere introduced in this studyto uncover the complex interplay of
ssing Technology 209 (2009) 4292–4311
different material and process parameters controlling the failure
mechanisms within the ASBs in cold heading process.
The ductile failure initiation under compressive loads is a
debated subject of many works in the literature. However, this
detailed study reached important findings that uncovered the
complex interplay of the different mechanisms of ductile failure
initiation and growth under compressive loads. All these findings
were based on detailed supporting experimental and metallurgical
evidences about these mechanisms.
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	Internal ductile failure mechanisms in steel cold heading process
	Introduction
	Cold heading process
	ASB phenomenon
	Examples of internal ductile failure in cold-headed products
	Experimental
	Material selection and characterization
	Testing material cleanliness
	Tests
	Sample preparation
	Results and discussion
	Void initiation mechanism
	Effect of material flow pattern on void nucleation
	Void nucleation mechanism at the inclusions-matrix interface
	Void nucleation mechanism at the grain boundaries and the second phase-matrix interface
	Ductile failure characterization in the testing materials
	Conclusion
	References