E J M Hiensch

J Beyer

P A Besselink

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199 
Resistance welding of TiNi-shape memory alloys 
J.Beyer, E.J.M. Hiensch, University Twente, Department of Mechanical 
Engineering, Laboratory for Materials Science, Enschede, M.L. 
P.A.Besselink, Foundation for Advanced Metals Science, (SCM) Hengelo, N.L. 
Introduction 
Joining processes as electronbeam welding, MID-welding or laserbeam welding, 
where the Joint is produced by melting of the material, are not suitable for 
shape memory TiNi-alloys due to the embrittlement of the welded region (1). 
Resistance butt-welding, especially of wires, has now been recognised as an 
effective process for Joining SMA-parts, while preserving the SHA properties 
in the welded region (2,3). However, large scatter of the resulting properties 
has been obtained in case the welding conditions are not controlled within 
narrow limits. Important variables are: heat input, compression force, 
alignment and shape of the contact surface of the wires (flat or pointed). 
A critical review of the welding process described in the previous paper (3) 
initiated the design and development of an improved welding instrument (4). An 
effort has been made to expand the practical utility of this process to a 
variety of wire diameters. The possibility of welding rings has been 
considered as a consequence of their large applicability. 
This paper reports the further development and results obtained by this 
improved experimental setup. 
Experimental 
A Ti-50.3 at.% Ni shape memory alloy with a diameter of 5.6 mm., as supplied 
by Krupp, has been taken as the starting material ( As-temperature of 313 K). 
The material was cold drawn to a variety of diameters (1.92, 1.43, 1.01, 0.49, 
0.19 mm). After every reduction of area of 25 - 30%, the material was annealed 
at 923 K, except for the last memory heat treatment, which was taken as 773 K 
for 5 minutes. From each diameter, pieces of 15 mm. length were taken for the 
welding experiments; from the wires with a diameter of 1.92 and 1.43 mm, one 
end was conically ground with an apex angle of 90-100 degrees. The contact 
area during welding was thereby 40% of the wire cross section for these 
diameters. The thinner wires, down to 0.19 mm, were welded with flat 
perpendicular surfaces. 
Determination of the transformation behaviour of base and weld material was 
carried out on a Stanton Redcroft DSC instrument. Tensile testing was carried 
out in a 20 kN tensile machine with a speed of 10 mm/min. Optical and electron 
microscopy (Jcol 200CX analytical TEM) were used to characterise the 
structural changes. TEM-specimens were prepared in a jet polishing machine 
containing a 7% solution of perchloric acid in methanol, at 285 K and 25 V. 
The welding instrument has been described in a previous paper (3). The working 
principal is as follows. A compressive force is exerted on the ends of two 
wires, fixed into one movable and one fixed clamp. A current, supplied by a DC 
power regulator, is led into the wires, thereby heating the contact area. The 
compressive force will press the weak material outward, thus making a sound 
weld. 
The following improvements have been made: 
- a more rigid construction in order to make the alignment more accurate and 
reproducable, which now occurs with an optical microscope 
- a whole set of clamping plates, one Tor each diameter (straight and rings). 
- the compressive force, exerted on the two parte to be welded, is made 
Article available at http://www.esomat.org or http://dx.doi.org/10.1051/esomat/198904003
200 
effectively constant during the welding cycle by using sets of disc-springs. 
instead of a single spring. 
In order to weld a larger variety of wire diameters, a DC power supply of Weld 
Equip with a max. output of 3600 A was used. Fig. la represents the complete 
set-up and fig. lb is a detail of the clamps with an as-welded ring. 
Fig. 1. 
a: An overview of the present welding instrument. 
b: Detail of the clamping system with an as-welded ring (wire diameter 2 am). 
201 
Results and discussion 
The optimum welding conditions, with respect to tensile strength, have been 
determined for each wire diameter. As an example, fig. 2 illustrates that 
there is a maximum tensile stress attainable for each series of compressive 
force exerted on the two parts to be welded. 
Tensile stress as a function of welding current for different compressive 
forces during welding. 
The lower strength at higher currents indicate fracture outside the weld, due 
to excessive heating at the clamps. Table 1 summarizes the maximum attainable 
tensile stresses relative to the tensile stress of the base material. The 
recoverable strain which has been measured by a bending test with a maximum 
strain on the circumference or the wire of 6%, is also given. 
Table 1. 
Diameter (ma) 1.02 1.43 1.01 0.49 0.19 
o: % of max. 
tensile stress 
8S 93 97 97 100 
shape recovery 
(X) 
87-100 97-100 97-100 100 100 
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These relative values are noticably higher than the previously reported ones 
(70-80%). Relatively lower strength levels of the larger diameter wires are 
observed. 
An explanation can be found in the conduction of the current from the clamps 
into the wire surface: as the wire diameter increases the current density at 
the surface will increase; the temperature at the surface is thus raised 
during welding. As a result, the structure will be weakened and failure 
outside the weld will occur on tensile testing. 
For thicker wires a reconsideration of the clamping mechanism will be 
necessary. In addition, to minimise the electrical resistance between clamps 
and wire, eventually leading to hot spots, it is necessary to remove the oxyde 
layer. 
It is important to notice, that the spread in the measured values was reduced 
to no more than 6%. This is mainly due to the better alignment possibilities, 
resulting in better parallelism of the wire contact surfaces. 
In order to obtain a proper weld, the Heat Input (HI) J/mm2 during the welding 
period (t1 + t 2 + t3) showed to be a dominant factor. 
HI = I2(t).R.dt and R = U/I 
where U = welding voltage and I = max.current (at t2) when U is reached. 
A linear relationship between HI and different diameters, as indicated in fig. 
3, is found for the wires up to 1 mm diameter, where a flat contact surface is 
used. The relationship only accounts for the initial length of the wire 
between the movable clamps. 
Fig. 3. 
a: Relation between heat input and wire diameter. 
b: The welding current is build up in 3 periods during welding. 
203 
The heat losses through the clamps and the change or 1 during welding are not 
considered. Experiments show, that if the change of length is limited, which 
can be done by variation of the welding time (t1 + t2 + t3) using the same 
heat input, the tensile stress of the weld can raise to a value even closer to 
that of the base material as indicated in table 1. The deviation of the 
calculated HI and the actual one for the results reported, showed to be within 
8X at present. The applied compressive stress and the current density can be 
calculated from the empirical relation 
r(d1)=r(d2). (d1/d2)1/2 and J(d1)=J(d2). ( d 1 / d 2 ) 1 / 2 
Although more data are required to be conclusive, the same type of behaviour 
is proposed for the larger diameter wires with conical shaped edges. Rings of 
the same material (ring diameter 21 mm) are also welded with the use of the 
same parameters. The same tensile strengths were obtained as reported for the 
straight wires (diameter 2 mm). 
The promising results of the tensile tests alone, are no guarantee for good 
fatigue properties. DSC measurements were carried out to elucidate the 
transformation behaviour of the weld relative to the base material. 
Fig. 4 shows a DSC curve for the weld as well as for the base (2 mm) wire. The 
heating and the cooling part of the curves have to be considered separately. 
On heating, the martensite - austenite transformation occurs at the same 
temperature. This is important, because the SKA effect occurs in this region. 
On cooling it can be noted, that the weld and the base material behave 
differently. The base material first transforms to R-phase (TR is 333 K) and 
the martensite is well separated from the R-phase with an Ms-temperature of 
Z71 K. The weld metal has a Ms-temperature of 285 K. These appreciable 
differences in transformation temperatures will lead to a build up of tensile 
stresses, either in the welded region or in the base material, depending on 
the relative position of the curves. 
x 
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Fig. 4. 
DSC-curves for the base material (a) and the weld (b). 
Successive transformation cycles can initiate early failure in the vicinity of 
the weld. This can than lead to the deterioration of the fatigue properties, 
while the SMA-part is in service. The DSC curve also indicates the lowest 
temperature (Mf) to which the material should be cooled in practice. The weld 
and base material should be fully martensitic before straining in order to 
obtain repeated SMA-behaviour. 
The optical and electron micrographs of fig. 5 show the microstructure of the 
weld and the base material. 
20S 
Fig. 5. 
a: Cross section of a 2 MS welded wire. 
b: Microstructure of the weld. 
c: TEM Micrograph of the austenitic base material, 
d: TEM micrograph of the martensitic weld 
The welded region in the optical micrograph is approximately 0.2 mm wide. The 
grain size in the weld is about 2 times smaller than in the base material. The 
weld itself seems to have less Ti2Ni particles than the heat affected zone, as 
was described previously (3). 
The electron micrographs show that the weld is transformed to martensite, 
whereas the base material is still austenitic. The electropolishing in this 
example is carried out at 283 K. 
The DSC results show that this temperature is Just between the two 
transformation regimes of the weld and the base material. On cooling this 
causes stresses resulting in the enhancement of the dislocation density. 
These results show that apart from the tensile testing, which reflects the 
optimum welding conditions, it is important to characterise the transformation 
behaviour by DSC (or other physical techniques). 
In order to test the usefulness of this welding technique, another alloy in 
the pseudo elastic condition has been welded. The first preliminary results 
indicate an even slightly, better tensile strength in the welds: 93% for 1.92 
mm wire and 97% for a 1.32 mm wire. It is speculated, that the electrical 
resistance of the B2 phase is higher than for the martensite, which leads to 
lower current density and less variation in properties due to hot spots at the 
clamps. 
Conclusions 
The newly designed welding instrument showed to produce wolds with SMA 
properties and behaviour which is comparable with that of the base material. 
An empirical relation is found for the heat input and the wire diameter, which 
makes the determination of the proper experimental parameters simple in 
practice. 
The applicability of this technique to straight and ring shaped material is 
demonstrated. 
Welding of martensitic as well as austenitic TiNi-alloys in various shapes is 
206 
possible and will be demonstrated in the near future. 
DSC measurements are a requirement, apart from mechanical testing, to specify 
the quality of the weld. 
(1) C.M. Jackson, H.J. Wagner and R.J. Wasilewskl, Nasa Report, SP 5110 
(1972) 
(2) N. Nishikawa, H. Tanaka, M. Kohda, T. Nagaura and K. Watanabe, Journal de 
Physique 43 (1982) C 4-839. 
(3) J. Beyer, P.A. Besselink and J.H. Lindenhovius, Proc SMA 1986 Conference, 
Cullin, China, 492. 
(4) N. van Leerdam, private communication, University Twente. 


Resistance welding of TiNi-shape memory alloys

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Resistance welding of TiNi-shape memory alloys

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