Science and Technology of Welding and Joining
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Dissimilar friction welding of NiTi shape memory
alloy and steel reinforcing bars for seismic
performance
Matthieu B. Lezaack, Aude Simar, Yannick Marchal, Martin Steinmetz, Koen
Faes & João Pacheco de Almeida
To cite this article: Matthieu B. Lezaack, Aude Simar, Yannick Marchal, Martin Steinmetz, Koen
Faes & João Pacheco de Almeida (2022) Dissimilar friction welding of NiTi shape memory alloy and
steel reinforcing bars for seismic performance, Science and Technology of Welding and Joining,
27:6, 418-428, DOI: 10.1080/13621718.2022.2061692
To link to this article: https://doi.org/10.1080/13621718.2022.2061692
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SCIENCE AND TECHNOLOGY OF WELDING AND JOINING
2022, VOL. 27, NO. 6, 418–428
https://doi.org/10.1080/13621718.2022.2061692
RESEARCH ARTICLE
Dissimilar friction welding of NiTi shape memory alloy and steel reinforcing
bars for seismic performance
Matthieu B. Lezaacka , Aude Simar,a , Yannick Marchala , Martin Steinmetzb , Koen Faesc and
João Pacheco de Almeidaa
a Materials and Civil Engineering (iMMC), Institute of Mechanics, Materials and Civil Engineering (iMMC), UCLouvain, Louvain-la-Neuve,
Belgium; b Architecture and Built Environment (LAB), Institute of Landscape, Architecture and Built Environment (LAB), UCLouvain,
Louvain-la-Neuve, Belgium; c Technologiepark-Zwijnaarde, Belgian Welding Institute (BWI), Technologiepark-Zwijnaarde, Gent, Belgium
ABSTRACT
ARTICLE HISTORY
Integration of NiTi superelastic bars with steel reinforcing bars is a promising path for increasing the seismic performance of reinforced concrete structures. Welding of NiTi alloy to steel is
however not straightforward with standard welding techniques. Friction welding is investigated
for performing this dissimilar junction using both regular friction (forced controlled) and milling
(position-controlled) machines. When adequate shortenings and forging forces are selected, the
welds can withstand the tensile strength of the steel rebars. Failure actually occurs in the heataffected zone of the steel bars. The intermetallic (IM) layer at the welded interface should be kept
thin and cracks should be limited in the steel layer next to the interface in order to avoid early
brittle interfacial cracking.
Received 16 December 2021
Revised 12 March 2022
Accepted 14 March 2022
Introduction
Reinforced concrete (RC) dominates the construction
sector internationally. A key contribution to addressing
the RC construction sector challenges, namely its carbon footprint, consists in designing structures for significantly longer service lives [1–3]. In seismic regions,
this implies higher probabilities of sustaining stronger
ground motions, which can cause excessive residual (or
permanent) displacements, as shown after the NewZealand earthquakes of 2010–2011 [4]. To minimise the
latter, harnessing the superelasticity of shape memory
alloys (SMA) [5] has recently become an active topic of
research[3, 6–14]. Since the response of RC structures
is largely governed by the reinforcing bars (or simply
rebars), the partial replacement of these ductile construction steel rebars – which depict large plastic strains
– by SMA reinforcements is an attractive and simple
solution to minimise residual structural displacements.
The efficacy of this approach has been proven from
investigations focusing on these regions of columns
[9,10], beam-column joints [11], beams [12], and
more recently from promising applications to RC walls
[3,13,14] with Nickel-Titanium (NiTi) SMA rebars.
One of the practical challenges for a more widespread engineering use of Nickel-based SMA rebars
is the connection between the latter and standard
deformed or ribbed steel rebars. Mechanical Nickelbased SMA-steel rebar connections are not desirable
CONTACT Aude Simar
KEYWORDS
dissimilar welding; friction
welding; shape memory
alloy; steel reinforcing bars;
force control; intermetallic
layer; tensile strength
for practical reasons as mechanical couplers are bulky,
which leads to many problems during seismic response.
A more fundamental connection between steel and
NiTi SMA rebars is required for leveraging the full
potential of NiTi-rebars reinforced structures: welding.
Akselsen et al. [15] describe the state of the art in
terms of welding shape memory alloys. They state that
classical techniques such as arc or laser welding can
affect drastically the SMA microstructure, by generating intermetallic (IM) phases or coarse dendrites. The
mechanical properties of these classical welds are thus
quite low, limiting the interest of welding SMAs [15].
Thus, it appears more appropriate to exploit the
potential of friction welding to perform reliable welds
between rod-shaped bars of SMAs. Friction welding is
a solid-state joining technique, reviewed by Maalekia
et al. [16] and Uday et al. [17]. Friction welding can
be processed by rotary, linear or orbital friction of the
components to be welded [16]. Mechanical energy is
transformed into heat at the interface of the parts. Thus,
energy losses are limited compared to fusion welding
techniques, which makes friction welding an efficient
process. Furthermore, the plastifying zone expells the
initial interface material, allowing to remove interface
oxides and impurities out of the weld core section [16].
Friction welding was successfully applied to materials
hard to weld such as high carbon content steels [16].
Unlike fusion welding techniques, no shielding gas or
aude.simar@uclouvain.be
Supplemental data for this article can be accessed here. https://doi.org/10.1080/13621718.2022.2061692
© 2022 Institute of Materials, Minerals and Mining. Published by Taylor & Francis on behalf of the Institute.
SCIENCE AND TECHNOLOGY OF WELDING AND JOINING
filler materials are required in friction welding. The
limiting applicability of the friction welding process is
related to the friction setup: at least one part should
move and the materials should deform plastically to
ensure a strong bonding [16].
Friction welding allows welding of dissimilar materials with rather highly efficient welds [18–20]. In particular, dissimilar welding is reviewed by Udey et al. [17].
Friction welding allows joining materials with highly
different mechanical and thermal properties, which is
often not successfull with classical fusion welding techniques [17]. Indeed, fusion welding techniques generate large and brittle intermetallic layers at the junction
of dissimilar materials, or can lead to solidification
defects, which are both avoided by the compressive
forces during friction welding [17].
Concerning dissimilar friction welding of NiTi
alloys, Fukumoto et al. [18] welded NiTi SMA alloy to
stainless steel via friction welding. However, in NiTi
SMA-steel friction welded samples, detrimental intermetallic layers of Fe2 Ti form during the process, leading
to a brittle behaviour of the joint [18]. The intermetallic layer can be modified by adding Ni sheets between
the rods before welding. This changed the intermetallic
composition to TiNi3 at the NiTi – Ni interface while
no reaction layer was observed at the Ni-steel interface. The resulting junction increased the weld tensile
strength up to 500 MPa, 150% higher than the Ni-free
interface version [18]. However, the addition of a Ni
interlayer makes the process more complex, expensive and labour intensive, which is not suitable for our
targeted application.
Owing to the past limited investigations on welding
SMA to steel using friction welding, the present work
will propose a detailed analysis of NiTi – steel welds
performed with this technique. One of the issues with
dedicated friction welding machines is their limited
availability for common civil engineering applications.
Indeed, conventional milling machines are available in
all machining workshops. The applicability of performing such welds on position-controlled machines is thus
also investigated. Using suitable parameters allows the
manufacturing of full-strength joints on both kinds of
setups, opening the path towards widespread applications for minimal residual displacements in structures
after earthquakes.
Experimental methods
Welding setups
The individual steel rebar units are nominally of Class
B according to Annex C of Eurocode 2 [21]. The nominal diameter is 12 mm, which is the equivalent plain bar
diameter having the same weight per unit length as the
deformed bar. The NiTi bars have a smooth surface and
a diameter of 12.7 mm. The steel and NiTi bars are cut
419
in 120 mm length parts to obtain the test units. Friction
welding is performed on two different machines:
(i) a dedicated friction welding machine (Thompson,
40 tons axial force machine) presenting a possibility to control the applied force during the forging
step (called ‘force control’, FC conditions);
(ii) a position-controlled (called PC conditions)
milling machine. This latest machine is actually
a dedicated friction stir welding (FSW) machine
(TRA-C industries) which, contrarily to a standard
milling machine, presents the advantage of having
a measurement of the forces during welding ideal
for research purposes. However, this machine did
not allow force control during the forging phase.
On the PC machine, the NiTi bar is inserted into
the rotating part while the steel rebar is clamped on
the machine table by a chuck. On the dedicated FC
machine, the NiTi part are inserted into the rotating
head of the machine and the steel on the stationary side.
The welding process is monitored in order to register
force, shortening displacement, time and revolutions
per minute (RPM).
Characterisation methods
Inductively Coupled Plasma Mass Spectrometry (ICP)
is used to determine the chemical composition of the
base materials. The latter, as well as the welds, are also
characterised by optical microscopy after standard polishing. Microstructures are revealed by etching the polished surfaces. Nital etchant (ethanol + 2% vol HNO3 )
is used for revealing steel microstructure, while NiTi is
etched by a solution of HF, HNO3 , and H2 O in volumetric proportions 1:2:10, respectively. Weld interfaces are
further characterised by Scanning Electron Microscopy
(SEM) for higher magnifications using a Zeiss-Ultra55
equipment. Chemical analysis of the weld interface is
performed inside the SEM by Energy Dispersive XRay analysis (EDX). X-ray diffraction (XRD) was performed on base materials on a 5 × 10 mm⊃2 zone as
well as on a welded specimen interface in order to identify the crystallographic phases. X-rays are obtained by
Co Kα radiations, the diffraction lines were recorded
from 2θ = 20° to 125° with 0.02° steps.
The mechanical properties of the base materials
and welds are also characterised. Micro hardness tests
(HV0.5) were performed with an Emco-Test Durascan 5G machine. Weld cross-section interface hardness lines are performed with 250 μm spacing between
indents. Uniaxial tensile tests on welds were performed
with a MTS 100 kN machine. Steel and NiTi bars,
300 mm long, are directly tested without extra machining following ISO6892-1 standard test specifications
[22]. Selected welded samples (see Section 3.2) are
machined into cylinders 8 mm diameter before tensile
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M. B. LEZAACK ET AL.
Table 1. Chemical composition of base materials (steel rebars and NiTi) in wt-%.
Material
Steel
NiTi
Fe
Ni
Ti
C
Cu
Mn
Mo
P
Si
Zn
O
98.150
0.009
0.105
55.880
/
44.054
0.210
0.034
0.445
/
0.615
/
0.020
/
0.010
/
0.175
/
0.025
/
/
0.023
Figure 1. (a) Uniaxial tensile response of base materials, (b) initial microstructure of NiTi, austenite phase is identified, (c) initial
microstructure of steel: ferrite and pearlite phases are identified.
testing. Broken samples are observed in SEM for fracture surface analysis.
Results
Base materials
The chemical compositions of the base materials,
obtained by ICP, are displayed in Table 1. Their mean
hardness values are 235 ± 9 and 237 ± 11 HV0.5 for
the steel and NiTi base alloys respectively. Figure 1(a)
provides the tensile curves of the base materials, highlighting the pseudo-elastic plateau of the NiTi material
at around 430 MPa. Steel has a higher yield stress of
530 MPa but a limited strain hardening in comparison
with the NiTi. Indeed, NiTi reaches 770 MPa as ultimate
tensile stress upon phase transformation to martensite.
In Figure 1(b) the microstructure of the NiTi is displayed, revealing equiaxed grains identified as austenite
phase. Figure 1(c) reveals the steel microstructure. Ferrite (note that the carbon content in Table 1 is 0.21
wt-%, hence inferior to the 0.76 wt-% eutectoid composition) and pearlite phases co-exist, and grains are
significantly smaller than NiTi grains: around 10 μm for
steel versus 50 μm for NiTi. Note that the steel rebars
contain coarse voids elongated in the milling direction,
up to 50 μm equivalent diameter.
Selection of the studied welding parameters
Welding parameters depend on the machine used.
While the PC machine controls the rotational speed
and the shortening displacement during welding, the
FC machine controls the rotational speed and the axial
force during welding. Welding trials are thus associated with a rotational speed, a maximum compressive
axial force (or welding force), a total shortening (friction + forging shortening cumulated) and a total friction time. The interest of evaluating the welding feasibility on a position-controlled (PC) machine is the
availability of such equipment for common civil engineering applications.
Almost 25 welding tests have been performed, whose
results are summarised in Figure 2 and in the supplementary material (Figure A.1). Figure 2 displays a
cartography of evaluated parameters in terms of total
welding shortening and welding force. The weldability window is identified as the welding conditions for
which the welded joint maximum tensile stress is larger
than the yield stress of the steel rebars ( > 530 MPa).
This corresponds to the structural engineering design
goal of enabling a pseudo-elastic cyclic response of the
shape memory alloy rebars. This condition minimises
the residual displacements of reinforced concrete (RC)
members, before plasticity spreads to the adjoining steel
rebar resulting in increasing member permanent deformations. It is interesting to note that a maximum force
of about 50 kN is reached in this weldability window,
requiring milling machines capable of holding such a
large compressive force.
For the following investigations, only four samples will be considered for illustrative purposes: two
performed on the PC machine and two performed on
SCIENCE AND TECHNOLOGY OF WELDING AND JOINING
421
Figure 2. Friction welding trials performed for this study. PC and FC setups are specifically identified. A weldability window was
identified at high force, as the tensile strength of the obtained welds is significantly larger than what was obtained at lower welding
forces. Four specific samples are discussed in this work, identified by black diamond. One of these samples is voluntarily outside the
weldability window.
the FC machine (indicated also in Figure 2). Their
appearance just after welding is provided in supplementary Figure A.2.
• A ‘PC brittle’ sample, outside of the weldability
region, representative of brittle behaviour of the
weld;
• A ‘PC necking’ sample, at the maximum investigated shortening, representative of ductile behaviour
of the weld and strain localisation followed by failure
in the steel part of the weld;
• A ‘FC necking’ sample, again representing ductile
behaviour and strain localisation followed by failure
in the steel part of the weld.
• A ‘FC no necking’ sample, characterising ductile
behaviour with fracture at the welded interface.
between Figure 3(c,d). The ‘FC necking’ sample of
Figure 3(d) has however a longer dwell time before
releasing the force after welding. A consequence of
the position control (PC) setup is that the compressive
force attains peaks, corresponding to increasing shortening, followed by significant force relaxation. Indeed,
the thermal retraction of the weld due to cooling is lowering the compressive force. The PC setup increases
the shortening in two steps because of the sequential
control that needs stopping the rotational speed before
activating material forging. The maximum peak force
of the ‘PC necking’ sample is around 50 kN, only slightly
lower than the 54 kN force that the ‘FC necking’ sample
attains at the end of the forging step. Comparing both
‘necking’ samples, the total shortening for the PC setup
doubles with respect to the FC setup sample (for similar
maximum applied force).
Detailed welding conditions
Figure 3 displays the processing curves associated to
the welding of the four representative samples. FC samples are associated to classical friction welding curves,
as shown in [16], while PC samples have rather unusual
force evolution for friction welding. It shows that the
‘PC brittle’ sample reaches a low maximum welding
force (26.9 kN) and limited shortening. In Figure 3(b),
the ‘PC necking’ sample attains higher welding force
and also higher shortening. In comparison, the FC samples in Figure 3(c,d) are subjected to higher forces,
around 55 kN, which are sustained during the following time interval, corresponding to classical friction
welding curves [16]. There is no significant difference
in terms of friction time, axial force and shortening
Microstructural characterisation of the welded
joints
Figure 4 shows the microstructure of the four investigated welds observed with scanning electron microscopy (SEM) at the weld centre. In Figure 4(a), the brittle
joint interface is highlighted. The ∼ 50 μm thick interface is composed of various intermetallic (IM) layers,
cracks and porous zones. The coarse IM layer thickness
of Figure 4(a) (weld centre) is not constant throughout the radial direction, as also observed by [18]. In
Figure 4(b), the ‘PC necking’ sample is showing a much
thinner interface. A few cracks aligned in the radial
direction, resulting of high deformations under welding, are still visible on the steel side. In Figure 4(c,d),
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M. B. LEZAACK ET AL.
Figure 3. Welding curves: compressive axial force, shortening and RPM of the bars during friction welding; (a) ‘PC brittle’ sample,
(b) ‘PC necking’ sample, (c) ‘FC no necking’ sample, (d) ‘FC necking’ sample. Note that the ‘FC necking’ sample has a holding time
twice longer than sample ‘FC no necking’ after forging, i.e. the welding force is maintained constant during cooling of the weld. On
all samples, the friction time is highlighted, defined as RPM > 0 and Force > 0 simultaneously. Friction time is longer for PC samples
(40 s and 50 s, respectively) than for FC samples (10 s in both cases).
Figure 4. (a) ‘PC brittle’ weld; a 50 µm thick reaction layer composed of intermetallic phases (IM) is formed at the interface, (b) ‘PC
necking’ sample; the reaction layer is reduced to less than 1 µm, (c) ‘FC no necking’ sample; the reaction layer is small but defects
(cracks and porosities) are visible at interface vicinity, (d) ductile ‘FC necking’ sample; reaction layer is even lower than 1 µm. Cracks
and coarse porosities are identified by white arrows on steel parts of (b) (c) and (d).
the FC samples have also thin interfaces, with similar
aspect. The thin IM layers seem homogeneous along
the radial direction at the scale of these SEM measurements, keeping a sub-micron size. Porosities and cracks
are observed on the steel part of ‘PC necking’ and FC
samples. Note that the interface of Figure 4(a) will be
further analysed via EDX, see Figure 7. Other samples
present too thin IM layers to be observable with EDX.
Figure 5 shows a detailed SEM characterisation of
the three interfaces corresponding to ductile failure.
SCIENCE AND TECHNOLOGY OF WELDING AND JOINING
423
Figure 5. Detailed interface microstructure: (a) PC necking sample, showing refined steel grains around the interface compared
to initial microstructure and martensitic NiTi structure (b) FC no necking sample, presenting finer steel grains and martensitic NiTi
structure, (c) FC necking sample, the finest steel grains are observed near the interface, a refined austenitic microstructure is visible
3 µm away from the weld interface as a martensitic layer had formed next to the interface. The intermetallic layer is thinner than 1 µm
in all conditions.
Microstructures of the two FC samples are not equivalent despite similar welding conditions in terms of force
and friction time in Figure 5(b,c). On the ‘FC necking’
weld, NiTi grains are austenitic but refined compared
to the initial grain size (see Figure 1(b)) near the weld
interface. For that sample, a thin martensitic layer of
3 μm has formed at the junction, oppositely to the ‘PC
necking’ and ‘FC no necking’ samples, where martensitic layers develop along tens of micrometres thickness.
This martensitic layer was indeed identified by XRD as
will be shown next on the sample ‘PC necking’ presenting the thickest layer. Concerning the steel part of the
weld, equiaxed and refined grains are observed compared to the initial structure (see Figure 1(c)). Note that
steel grains are larger for the ‘PC necking’ sample.
Figure 6 shows the XRD analysis of base materials and weld interface. The peaks for the steel material correspond to ferrite and body-centred cubic BCC
microstructures in accordance with Zhang et al. [23].
Peaks in NiTi material correspond to austenitic structure, as shown by Dudr et al. [24] and Fukumoto et al.
[18]. The weld interface XRD measurement reveals that
a new phase had formed after friction welding, as a
new rising peak is visible. This new peak is related to
Martensitic NiTi, as observed by Dudr et al. [24].
Figure 7 shows the EDX line analysis of the chemical
composition of the intermetallic layer of the ‘PC brittle’
sample. The approximatively 40 μm thick interface is a
mix of the main chemical elements: Ni, Ti, and Fe in
varying proportions depending on the position. On the
steel side, two or three Fe-rich intermetallics are generated with various Fe content. On the NiTi side, a highly
porous IM layer presents a reduced Fe content.
Mechanical characterisation of the selected
welded joints
Figure 8 displays the hardness of the weld mid crosssection for all samples. Figure 8(a) displays the hardness
profile at a 10 mm distance from the weld interface on
both sides, while Figure 8(b) highlights the interface
Figure 6. DRX results on base materials and weld interface of
‘PC necking’ sample; steel base alloy is identified as ferrite and
body-centred cubic (BCC) Fe + C phases, NiTi base alloy is identified as austenitic NiTi, while the welded interface shows a rising
peak corresponding to martensitic NiTi.
hardness variation. Left side of Figure 8(a) reveals a
softening of the steel, reaching a minimum hardness
at 4 mm (‘PC brittle’) or 8 mm away from the interface (all other samples). Right side of Figure 8(a) shows
that NiTi hardness has a minimum value at 3 mm away
from the interface for both ‘FC’ samples before a slight
increase up to ∼ 275 HV0.5, 10 mm away from the
interface. The ‘PC’ samples NiTi side have more homogeneous hardness, around 230 HV0.5 (which is the
initial hardness of the material). The full weld section
hardness map of ‘FC necking’ sample is available in
supplementary Figure A.3.
In Figure 8(b), the hardness near interface region
is highlighted. Steel hardness is increasing when
approaching the interface, as does the NiTi hardness
(except for the ‘PC brittle’ sample that shows homogeneous 240 HV0.5 hardness). Local peak hardness near
the weld interface is not related to IM layer for ‘FC’s
and ‘PC necking’ samples, as the indents are at least
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M. B. LEZAACK ET AL.
Figure 7. EDX measurements on ‘PC brittle’ sample interface. A 40 µm reaction layer generates intermetallic phases containing Fe,
Ni and Ti in varying proportions.
Figure 8. (a) Overview of the mid-section hardness profile through the four studied welds, (b) zoom on the interface region. Indent
spacing is 250 µm.
50 μm away from the IM layer. However, the proximity of the IM layer could constrain plasticity during
indentation and cause the appearance of an additional
parasitic hardening effect. Oppositely, the peak value at
interface for the ‘PC brittle’ sample is related to partial
indentation in the thick ( ∼ 40 μm) IM layer.
Figure 9 reveals the macroscopic tensile mechanical behaviour of the four selected welds. While the
‘PC necking’ sample has a ductile deformation, leading
to 9% engineer fracture strain, the ‘PC brittle’ sample and all similar samples of the non-weldable zone
of Figure 2 break under elastic deformation regime,
at extremely low stresses and strains ( < 200 MPa and
< 0.5%, respectively). The two samples obtained with
the FC machine show a ductile tensile response. The
‘FC necking’ sample has a similar deformation capacity
compared to the ‘PC necking’ sample, whereas the ‘FS
no necking’ sample breaks earlier without significant
necking, which is apparent from the absence of a final
unloading branch. Both necking samples failed in the
steel part of the weld outside the central 25 mm gauged
by the extensometer, while the ‘PC brittle’ and the ‘FS
no necking’ samples broke at the welded interface. Base
material curves are provided for comparison purposes,
evidencing the attainment of the ultimate steel rebar
strength, as originally desired.
In Figure 10, the broken surfaces of the four samples are investigated. In Figure 10(a), the weld interface
failed, revealing a flat surface composed of the intermetallic layer (Fe-Ni-Ti). Figure 10(b) reveals the broken surface of sample ‘FC no necking’, showing the
influence of the intermetallic layer on fracture path.
In that case, the intermetallic layer is more dense than
what can be observed in Figure 10(a) and leads to a
flat decohesion surface. In some regions of the fracture
surface of the ‘FC no necking’ sample the steel material is reached, as proven by Supplementary Figure A.4.
In Figure 10(c), the ‘FC necking’ sample and ‘PC necking’ both reveal a very similar fracture surface: a ductile
failure in the steel part of the weld. Ductile damage is
visible, as a wide population of cavities and porosities
are visible on the fracture surface.
Discussion
Effect of the welding parameters/conditions on the
joint quality
Many welding trials were performed on both PC and
FC machines (see Figure 2 and supplementary Figure
A.1). The PC machine allowed the control of the shortening at the friction stage (i.e. while the parts are
SCIENCE AND TECHNOLOGY OF WELDING AND JOINING
425
Figure 9. Tensile response of ‘PC brittle’, ‘PC necking’, ‘FC necking’ and ‘FC no necking’ welded samples in contrast with the base
material tensile curves; ‘PC brittle’ sample breaks at low stress and low strain, in the elastic regime. Oppositely, both ‘PC necking’ and
‘FC necking’ samples show the same behaviour and fail by necking at a strain corresponding to the crossing of the base materials
tensile curves, around 9% strain.
Figure 10. (a) ‘PC brittle’ sample failure surface, porosities are observed, (b) ‘FC no necking’ sample: fails at the interface but no more
porosities on fracture surface, some regions of the crack surface reach the steel material (see Supplementary Fig A.4), (c) ‘FC necking’
or ‘PC necking’ sample fracture surfaces, showing ductile damage by void nucleation growth and coalescence.
rotating) and at the forging stage. The very first trials were performed at limited shortening during the
friction stage, without forging shortening. The resulting welds were brittle and resulting tensile strength was
lower than 40 MPa. Iterations were done in order to
increase progressively the total shortening, by adjusting
the friction shortening as well as the forging shortening
(see supplementary Figure A.1). Below 20 mm of total
shortening, none of the obtained welds were successful
in rising the strength above 150 MPa. The examination
of these welds after fracture revealed a thick intermetallic layer, of the same nature than the displayed interface
in Figure 4(a). The thick intermetallic layer is inhomogeneous along the radial direction, in terms of existence
of several IM compositions and distributions, as well as
in terms of total IM layer thickness.
It was observed that limited shortening of the bars
is associated to lower compressive forces during the
forging steps on the PC machine, as seen for the
‘PC brittle’ sample in Figure 3. Low shortening and
low forces result in thicker intermetallic layers. This
layer is detrimental to the mechanical properties of the
weld as also demonstrated by Fukumoto et al. [18].
The intermetallic layer is a preferential path for crack
propagation as it already contains cracks after forging
(Figure 4(a)) as well as porosities. The intermetallic
zone is mainly composed of Fe2 Ti brittle phase and
TiNi3 phase, according to Fukumoto [18], which is in
good agreement with the results of Figure 7, as the Ni,
Ti, and Fe are coexisting inside the reaction layer.
Increasing drastically the shortening improves the
mechanical properties of the weld on the PC setup, as
proved by the ‘PC necking’ sample, see Figure 9. In
that case, the brittle intermetallic layer is driven out of
the interface during forging [16], following metal plastic deformation. Expelled material generates a curved
flash around the welded section (see Supplementary
Figure A.3). Larger shortenings necessarily increase the
amount of expelled material. The significant reduction of this interface intermetallic layer due to large
plastic deformations was not observed by Fukumoto
et al. [18], who obtained larger intermetallic layers in
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M. B. LEZAACK ET AL.
their reference welds, i.e. without using an interlayer Ni
sheet. Furthermore, the intermetallic layer at the interface obtained in the current welds inside the weldability
window is homogeneous in the radial direction, which
contrasts with the results by Fukumoto et al. [18], where
significant thickness variations were reported along the
radial direction.
Interestingly, the total shortening of ‘FC’ samples
during friction welding is about half of the ‘PC necking’ shortening (Figure 2) with equivalent peak force
(around 50 kN). Friction time is defined as the duration of the processes under which the force is positive (Force > 0 kN) and the parts are in rotation
(RPM > 0). Figure 3 suggests that, provided that sufficient heat is generated during the friction stage, similarly high weld strengths can be produced. As it can
be seen for the samples ‘PC necking’ and ‘FC necking’, this heat can be produced through longer friction
phases at lower compressive forces (‘PC necking’), or
shorter friction phases at higher compressive forces
(‘FC necking’). According to Figure 3, the friction time
of FC samples is close to 10 s at high compressive
force (15–20 kN), while the PC samples have approximately 40–50 s friction time at much lower compressive
forces. The RPM is also higher on the PC machine,
1500 rev min−1 oppositely to the FC machine that
worked under 1250 rev min−1 . The increased duration
and higher RPM could heat up the material to higher
temperatures at the vicinity of the weld interface despite
the lower compressive forces. This could facilitate the
deformation of the interface region during forging for
the PC samples, leading to larger deformation at lower
compressive force.
Effect of the welding conditions on the
microstructure
The plastic deformation of the interface during welding (mostly in the forging step) has a strong influence
on the initial materials’ microstructures. It is obvious in Figure 5 that steel grains (of all samples) are
much smaller than the initial grains of Figure 1(b). The
heat and plastic deformations led to dynamic recrystallisation (DRX) of the grains near the interface, as
observed by Satyanarayana et al. [20] on steel to steel
friction welded joint. Furthermore, the steel part of the
weld contains cracks and porosities, resulting from the
intense deformation (indicated by the white arrows in
Figure 4). Cracks may have formed by the collapse of
initial porosities contained in the steel bar, in addition to nucleation of new cracks by intense shearing
in the radial direction. Despite these relatively coarse
defects ( ∼ 10 μm long, and many smaller sizes), failure occurs 20 mm away from the interface on the steel
side of the ‘FC necking’ sample, (see Supplementary
Figure A.5). Oppositely, the ‘FC no necking’ sample is
probably affected by the presence of these coarse defects
as the fracture surface (Figure 10(b)) reaches the pure
steel material in some regions, as demonstrated in Supplementary Figure A.4. A critical size of defect could
thus explain the difference between the ‘FC no necking’ and the ‘FC necking’ samples. Longer dwell time
after forging could be related to better closure of the
cracks near the interface in the ‘FC necking’ welding
conditions.
Similarly to the steel part, NiTi initial austenitic
grains appear to be significantly refined in the sample ‘FC necking’ of Figure 5. It was shown that Tisteel joints caused DRX in the Ti part of the weld, as
described by Dey et al. [19]. The mechanism is thus
related to this recrystallisation under high temperatures and strains. It was also demonstrated by Mani
Prabu et al. [25] that high temperatures and strains
initiate dynamic recrystallisation when welding NiTi
plates via friction stir welding. According to the properties of NiTi alloy, high temperature and high strains
play opposite roles in the microstructure crystallography. High temperature favours austenite formation,
while high strain favours the formation of detwinned
martensite, as described by Xie et al. [26] and Jacobus
et al. [27]. Furthermore, strain rates influence phase
transformations: it was shown by Xie et al. [26] that the
stress needed for austenite to martensite transformation
is higher under higher strain rates. Related to our work,
the strain rates obtained during friction welding could
rise the austenite to martensite transformation stress to
higher levels, reducing the final amount of martensite
near the interface, even if this aspect was not further
investigated.
Focusing again on the weld manufacturing, friction
welding induces first high temperatures before applying large strains during the forging step. The NiTi
material close to the interface is thus austenitic before
welding. Increasing the temperature is not affecting
the austenitic nature of the NiTi but once the forging
starts, the NiTi layer close to the interface turns into
detwinned martensite phase under high stresses. The
NiTi martensitic zone that develops near the interface
is larger in the ‘PC necking’ and ‘FC no necking’ samples compared to the ‘FC necking’ sample, according
to Figure 4. The existence of the martensitic NiTi in
‘PC necking’ sample was shown by X-ray Diffraction
analysis in Figure 6.
The thinner martensitic zones of the ‘FC necking’ sample could be related to microstructural rearrangements under high pressures and temperatures, as
the holding time after forging is longer. As shown in
Figure 5(c), austenitic refined grains are visible near
the weld interface, finer than the initial NiTi grains.
Recrystallisation of the microstructure is thus obvious and could have transformed the martensite straininduced zones into austenitic grains, as observed in
other thermomechanical process at high strains and
temperatures [25].
SCIENCE AND TECHNOLOGY OF WELDING AND JOINING
Despite recrystallisation, a martensitic layer survives
at the interface vicinity, as displayed in Figure 5. The
martensitic layer thickness could explain the difference
between the ‘FC necking’ and ‘FC no necking’ samples
tensile behaviours (in addition to the defects in steel
part, as discussed above). Indeed, the larger martensitic layer of ‘FC no necking’ sample is associated to
significantly higher hardness at the weld interface (see
Figure 8), as referred with respect to other dissimilar welds by Uday et al. [17]. Oppositely, the thinner martensitic layer of ‘FC necking’ sample gives a
smoother hardness profile when crossing the interface
(see Figure 8). Reducing hardness gradients lowers the
strain localisation at the interface, which could prevent
brittle failure under straining.
Links between microstructure and mechanical
properties
According to the discussed effect of welding parameters on the microstructure, the highest compressive
forces during welding lead to the finest interface layers.
It is then proved on Figure 9 that the coarse intermetallic layer of sample ‘PC brittle’ results in poor
mechanical behaviour, breaking in the elastic regime
at 120 MPa. All other investigated PC samples that do
not reach high forces, i.e. above 30 kN (see Figure 2
and supplementary Figure A.1), show brittle behaviour
and limited tensile strength as the described ‘PC brittle’ sample. Oppositely, the thin intermetallic layers of
‘PC and FC necking’ samples allow the joints to reach
the ultimate tensile strength of the steel, as the interface does not break under loading, hence attaining the
original goal of this research. The effect of intermetallic
layers on tensile behaviour was investigated in the literature (Refs. [28, 29]). It is thus admitted that the tensile
stress rises with decreasing intermetallic layer thickness
under a specific size threshold. Full strength weld is thus
achieved, activating the pseudo-elasticity effect in the
NiTi bar, at least a few mm away from the weld interface
where the material is unaffected by the welding process
temperature and pressure.
However, despite similar intermetallic layer thickness (see Figure 5), the ‘FC no necking’ and the ‘FC
necking’ samples have different behaviour. The former
breaks at the interface, meaning that another feature
is affecting the fracture location. A first explanation
could be related to the existence of remaining defects
after welding close to the interface, as some cracks
are already visible in Figure 4(b–d). The presence of
a larger crack could increase the risk of interface failure, as shown in Supplementary Figure A.4, where the
crack surface crosses through the steel material in some
regions.
According to the hardness measurements in
Figure 8, the high temperatures and strains near the
welded interface affect the hardness of the material.
427
The weld interface has hardened in both NiTi and steel
vicinities compared to the initial hardness of both base
materials. This localised hardening is mainly explained
by the dynamic recrystallisation in steel regions and
martensitic transformation in the NiTi part of the weld.
However, steel hardness decreases within a minimum
8 mm from the weld interface, due to the progressive reduction of the amount of recrystallised grains.
This softening in the steel affects the strain localisation
under uniaxial loading and favours necking in steel part
of the weld, as observed for both ‘FC necking’ and ‘PC
necking’ samples.
Conclusions
The present work demonstrated that welding NiTi
bars to steel rebars is feasible. Full strength welds
were obtained using both a dedicated friction welding
machine (force-controlled FC setup) and a millingtype machine (position-controlled PC setup). Shortening during welding is the first order parameter in the
PC machine setup, allowing for high welding forces
and high strains. The PC machine has a longer friction
time, leading to extensive softening of the base materials around the weld interface. Such effect is limited
in the FC machine, allowing smaller shortenings for
identical peak forces.
In the analysed samples failing by necking, the interface between steel and NiTi is similar independently
of the machine used for welding. The thin interfaces,
less than 1 μm thick, are associated to a strong bonding and excellent tensile properties. In some cases, the
welded sample breaks after necking in the steel part of
the weld, while similar welding conditions in terms of
friction time, shortening, and forces, can break at lower
strains at the welded interface, without necking. This
could be explained by the presence of coarser defects in
the steel part of the ‘no necking’ welds and the smoother
hardness profile near the weld interface, limiting the
localised straining at the brittle interface layer. However, full strength welds can be achieved in both cases,
independently of the failure mode.
This welding process can be safely performed on
classical position-controlled milling machines. Under
loading of the composite NiTi-steel structure, the excellent bond will activate the pseudo-elasticity of the NiTi
parts, limiting plastic irreversible deformation. The
present work thus highlights the applicability of welding NiTi to steel rebars to improve the seismic response
of reinforced concrete structures. Future studies will
address the response of these welds under cyclic and
dynamic loading.
Acknowledgements
The authors are grateful to the technical staff of the LACAMI
and LEMSC technological platforms of iMMC. M.B.L.
acknowledges the support of a FRIA grant, Belgium. This
428
M. B. LEZAACK ET AL.
work was supported by the Fonds de la Recherche Scientifique – FNRS under Grant(s) n° F.4501.21.
[13]
Disclosure statement
No potential conflict of interest was reported by the author(s).
[14]
Funding
This work was supported by Fonds De La Recherche Scientifique - FNRS [grant number F.4501.21]; Fonds pour la Formation à la Recherche dans l’Industrie et dans l’Agriculture FRIA.
[15]
[16]
[17]
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