Min Invas Ther & Allied Technol 2000: 9(1) 107–118
`
`Optimisation of processing and
`properties of medical grade Nitinol
`wire
`
`A.R. Pelton1, J. DiCello1 and S. Miyazaki2
`
`1Cordis Corporation – Nitinol Devices and Components, Westinghouse Drive, Fremont CA, USA; and 2Institute
`of Materials Science, University of Tsukuba, Ibaraki, Japan
`
`Summary
`The purpose of this paper is to review the current processing and resultant properties of standard
`Nitinol wire for guide-wire applications. Optimised Ti-50.8at%Ni wire was manufactured according to
`industry standards by precise control of the composition, cold work and continuous strain-age
`annealing. Mechanical properties of this wire are reported from ⫺100°C to 200°C to demonstrate the
`effects of test temperature. Within the ‘superelastic window’ the plateau stresses are linearly related to test
`temperature. Additional ageing treatments can be used as a tool to fine-tune transformation temperatures and
`mechanical properties. A review of the fatigue properties of thermomechanically-treated Nitinol wire shows that
`they are affected by test temperature, stress and strain.
`
`Keywords
`Nitinol, shape-memory, superelasticity, mechanical properties, ageing, fatigue
`
`Introduction
`The growth in the use of Nitinol in the medical
`industries has exploded over the past 10 years.
`Patients and care-providers have encouraged the
`transition from traditional open-surgical procedures
`that require long hospital stays, to less-invasive
`techniques, which are often performed in out-patient
`facilities [1]. This demand for minimally-invasive proc-
`edures has required novel instrumentation and
`implants to be designed by engineers and physicians.
`An increasing number of these devices use Nitinol as
`the critical component. Examples of these medical
`applications are richly illustrated in companion articles
`in this journal [2,3], and range from endoscopic
`instruments to implants, such as stents and filters. It is
`interesting that the majority of these devices depend
`
`on mechanical superelastic behaviour, which is a
`significant departure from the original thermal shape-
`memory industrial uses of Nitinol.
`Since the ‘discovery’ of the shape-memory effect in
`TiNi alloys in the 1960s, metallurgists have investigated
`methods to control transformation temperatures and
`mechanical properties through alloying additions,
`improved melting practices and thermomechanical
`processing (see, for example, References 4 and 5).
`The production of many thousands of kilometers of
`wire for such diverse products as cellular telephone
`antennae, eyeglass frame components, guidewires,
`undergarment supports and orthodontic archwires
`profoundly influenced the acceptance of Nitinol in the
`marketplace. These commercial opportunities have
`allowed Nitinol suppliers to focus on improving
`
`Correspondence: A. Pelton, Cordis Corporation – Nitinol Devices and Components, 47533 Westinghouse Drive, Fremont CA 94539, USA.
`
`© 2000 Isis Medical Media Ltd
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`processes for a few standard alloys, rather than
`pursuing a myriad of ‘boutique alloys’ with niche
`applications. The composition and processes have
`been refined so that, for example, the transformation
`temperature in final products is routinely controlled
`to within ± 3°C. More recently, the availability of
`seamless tubing and sheet have provided designers
`with additional tools to solve engineering problems.
`Furthermore, microfabrication techniques, such as
`laser machining [6] and photoetching [7], have also
`contributed to the increase in the number of miniature
`devices made from Nitinol.
`Accordingly, Nitinol properties have become very
`predictable, which is a basic requirement of design
`engineers. As the Nitinol industry has matured over
`the past two decades, terms such as ‘shape-
`memory’, ‘superelasticity’, ‘recovery forces’, ‘plateau
`stresses’, and ‘transformation temperatures’ are now
`recognised by more
`than
`just a select
`few
`metallurgical specialists. Although design engineers
`have a good understanding of the basic properties of
`the alloys, they still have many good questions.
`Typically these include:
`
`■ Are the mechanical properties constant over a
`wide range of temperatures?
`■ Can we adjust the transformation temperature
`without modifying the mechanical properties?
`■ Do the shape-memory and superelastic properties
`imply that Nitinol has an infinite fatigue life?
`
`The purpose of this article is to address these
`questions by reviewing the processing and resultant
`properties of Ti-50.8at%Ni wire that has been
`manufactured for medical guide-wire applications.
`Furthermore, this article will focus on the effects of
`standard continuous thermomechanical processes,
`rather than long-term ‘batch’ processing, which was
`more common in the 1970s and 80s.
`Processing
`Optimisation of the superelastic properties of Nitinol
`for a specific product
`is achieved through a
`combination of cold work and heat treatment. The
`first step
`in optimising
`the
`thermomechanical
`treatments of wire and tubing products is to draw the
`material through a series of dies, to achieve 30–50%
`reduction in cross-sectional area [8]. Past methods
`have employed a long-term batch annealing process
`but, to attain a more uniform product, continuous
`strand strain annealing is the most effective method.
`With this method, the Nitinol wire is under constant
`strain during the annealing process. Continuous
`‘strain annealing’ ensures that the entire spool will be
`processed with the identical thermomechanical
`
`treatment, resulting in a product with uniform pro-
`perties from end to end.
`Figure 1 shows a typical continuous-strand strain
`straightening process line.
`Continuous-strand straightening usually occurs in
`a temperature range of 450–550°C under a stress of
`35–100 MPa. As the wire moves into the heat zone it
`will initially want to shrink in length and grow in
`diameter, due to the shape-memory effect not
`suppressed by the cold work of the drawing process
`(springback). The wire temperature quickly increases,
`its strength will drop and the applied strain will
`straighten the wire and, depending on the strain,
`reduce the wire diameter slightly. In the continuous
`process it is difficult to measure the active strains
`during straightening as they occur inside the furnace.
`However, the following vertical straightening example
`can help define the strains operating during the
`continuous process.
`During vertical straightening of discrete lengths of
`wire, the wire is heated via electrical resistance and is
`therefore exposed to make visible measurements. As
`a 1.5 mm diameter wire was being electrically
`straightened, it showed a maximum springback of
`1.2% during heating and a 2.4% extension strain at
`the end of the straightening cycle. Similar strains
`would be expected with the continuous-strain
`straightening method.
`Straightness, mechanical properties and the active
`Af are all affected by the speed and temperature
`parameters of the straightening process. As with all
`thermally-activated processes, time at temperature
`controls the final properties of the wire. More time at
`temperature softens
`the wire and moves
`its
`mechanical properties toward a
`fully annealed
`product, while short times leave the material closer to
`the high strength cold-worked state. To optimise the
`superelastic properties, a balance must be developed
`between these two extremes. The requirements of
`
`Figure 1. Schematic diagram of a continuous strand
`annealing equipment for optimised production of Nitinol
`superelastic wire.
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`Optimisation of processing and properties of medical grade Nitinol wire
`
`the final product will help define the process
`parameters. A high torqueable guide-wire may
`require slower speeds than a high strength wire that
`has a table-roll straightness requirement. Additional
`speed and
`temperature adjustments may be
`necessary to meet any active Af requirements (see
`discussion below) as well.
`
`Properties
`the methods of
`In this section we will consider
`characterising the thermal and mechanical properties
`of thermomechanically-processed wire. The focus will
`be on products that are superelastic between room
`temperature and body temperature. We will document
`the mechanical properties from ⫺100°C to 200°C, to
`illustrate how test temperature affects performance.
`Furthermore, since many wire and tubing products are
`given additional
`thermal shape-setting, we will
`establish the effects of ageing time and temperature
`treatments on transformation and mechanical
`properties. Finally, since many Nitinol medical devices
`are used in (human) fatigue environments, we will
`discuss the influence of strain amplitude and test
`temperature on bending fatigue behaviour.
`
`Transformation temperatures
`Harrison [9] cataloged >20 techniques that have been
`used to measure the changes associated with the
`shape-memory transformation. Two of the most
`fundamental ways to determine the qualitative
`transformation temperature involve sound and feel.
`Even a novice can distinguish between the ‘ping’ from
`austenite and ‘thud’ from martensite when dropped
`on the floor. Although this is not a highly quantitative
`means to measure transformation temperatures, it
`has been used to sort alloys quickly without
`sophisticated equipment. Another qualitative method
`is to feel the alloy. Martensite ‘feels’ rubbery when
`bent, whereas austenite feels ‘springy’.
`These two simple examples cited above illustrate
`that the martensitic transformation affects a variety of
`properties. However, Harrison [9] offered the sage
`advice that
`the chosen measurement
`technique
`should parallel the actual function of the product.
`For example, many medical customers request
`certification of the Af temperature to ensure that the
`product is austenitic above a certain application
`temperature. These customers generally specify that
`the measurements are obtained by either differential
`scanning calorimetry (DSC) or free recovery (‘active’)
`techniques. DSC measures the heat released and
`absorbed during the martensitic (exothermic) and
`austenitic (endothermic) transformations, respectively
`[10, 11]. Free recovery, however, is by far the most
`
`simple and often the most useful method to measure
`Af . This technique only requires the following steps,
`which simulate a shape-memory cycle:
`
`■ cool to a low temperature (for example in a cooled
`alcohol bath);
`■ bend the sample to a prescribed strain (2–3%);
`■ watch and record the temperature at which the
`sample returns to its original shape when heated:
`this is defined as the Af temperature.
`Free recovery can also be instrumented in order to
`obtain a permanent record of the results [12, 13].
`Both of the above techniques have the benefit that
`they are straightforward to conduct, amenable to use
`for small specimens,
`require minimal sample
`preparation
`(especially
`free
`recovery) and are
`reproducible.
`Figure 2 compares the DSC thermogram and
`instrumented free-recovery measurements from the
`same wire. The DSC records the heat flow during
`both cooling and heating, whereas free recovery
`records the deflection recovery only during heating.
`The key transformation temperatures, martensite start
`(Ms), martensite finish (Mf), austenite start (As) and
`austenite finish (Af), are marked as appropriate on
`both charts. Also included are atomic models of the
`austenite (cubic structure) and martensite (monoclinic
`structure)
`to help
`the
`reader visualise
`the
`transformations.
`Note that the DSC graph shows an R-phase peak
`during cooling from a high temperature. Although an
`in-depth discussion of the R-phase is beyond the
`scope of this paper, it is important to point out that
`the R-phase is another shear transition in competition
`
`Rs
`Rf
`
`Martensite
`
`50
`
`Figure 2. Differential scanning calorimetry and free recovery
`of the same processed wire. Note that upon cooling the wire
`transforms to R-phase prior to the martensitic
`transformation. Upon heating, both techniques provide
`similar As and Af temperatures as the (monoclinic) martensite
`transforms to the (cubic) austenite.
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`with martensite [14]. In the case shown in Figure 2,
`the R-phase forms around room temperature. With
`further cooling, the martensite transformation begins
`at about ⫺23°C (Ms) and is fully martensitic below
`⫺38°C (Mf). When the sample is reheated, the reverse
`transformation begins at about 22°C (As) and finishes
`at an Af of 32°C. Note that there is a wide hysteresis
`between the Mf and Af, which is characteristic of
`shape-memory alloys. The origin of the hysteresis is
`attributed to microscopic internal friction effects [15].
`(32°C) data from the free
`The As (24°C) and Af
`recovery method are nearly identical to those from
`DSC.
`
`Mechanical properties
`Based on the transformation behaviour shown in
`Figure 2, the wire should be fully austenitic above
`32°C. However, an important characteristic of shape-
`memory alloys
`is that stress can trigger the
`martensitic transformation at temperatures above Af
`(the ‘thermoelastic’ effect [15]). From a thermo-
`dynamic viewpoint, this means that it is easier (lower
`free energy) for the wire to create martensite in
`response to the applied stress than to deform
`plastically (dislocation formation) [15,16]. Stress-
`induced martensitic transformations can be easily
`understood by considering Figure 3 [17]. This dia-
`gram compares atomic motions in response to an
`applied stress by traditional Hookian elastic (top) and
`
`transformational superelastic (bottom). For Hookian
`elasticity, which represents conventional materials,
`such as stainless steel, the atomic bonds ‘stretch’ up
`to about 0.5% before plasticity occurs. In contrast,
`the austenitic structure depicted on the bottom left
`structure transforms into martensite with applied
`stress. As the magnitude of the stress increases from
`left to right, the amount of martensite increases. Up to
`10% strain can be accommodated by stress-induced
`martensitic transformations. The martensitic structure
`formed by superelasticity is identical to that formed
`through the shape-memory process, as illustrated in
`Figure 2.
`Figure 4 is a schematic stress–strain curve that
`corresponds to this model of transformational
`superelasticity. When the wire is pulled beyond its
`Hookian elastic limit (approximately 1.5% strain in
`Nitinol), there is an apparent ‘yield’ at a critical stress.
`Atomistically, this is represented by the onset of
`martensitic transformation, as shown in the second
`diagram in Figure 3. The wire can be further stretched
`at a relatively constant stress along the ‘loading
`plateau’ until the entire structure has transformed into
`martensite. As the stress is removed, the martensite
`immediately recovers elastically (linear unloading) and
`then begins to revert back to austenite on the
`‘unloading plateau’. The ability of the material to
`return to its original shape when the stress is
`released is an important attribute for many products,
`
`Figure 3. Schematic representation of the atomic motions associated with Hookian elasticity observed in conventional
`materials and transformational superelasticity of Nitinol. (From Stöckel and Yu, Ref. 17.)
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`at 100°C, with a high permanent set, but a well-
`defined unloading curve, indicates that deformation is
`accommodated by a combination of stress-induced
`martensite and conventional plasticity. Above 150°C,
`however, the wire deforms by plastic mechanisms
`rather than martensitic transformations, which results
`in a linear unloading curve. The temperature where it
`is too difficult to stress-induce martensite is defined as
`Md; in the present case, Md is between 100°C and
`150°C.
`The effects of test temperature on the tensile
`curves shown in Figure 5 may be further analysed by
`considering each of the key attributes. For example,
`Figure 6 shows the temperature dependence of the
`permanent set from these wires after unloading from
`6% strain. At lower temperatures, the unresolved
`strain is due to deformation of the martensite, and
`can be recovered by heating above Af. The residual
`strain is nearly zero between 0°C and 60°C, which
`defines the superelastic ‘window’ for this alloy. As
`noted above, the non-recoverable plastic strain is
`about 1 % at 100°C and then increases to about
`3 % at 150°C. Many medical applications require
`superelastic behaviour between room temperature
`and body temperature. Therefore, this 60°C window
`is perfectly centered about the intended application
`range.
`Figure 7 shows the effects of test temperature on
`the loading, unloading and ultimate tensile stress. We
`see that there is a linear relationship between plateau
`stress and temperature between about 0°C and 60°C
`for the unloading plateau and up to 150°C for the
`loading plateau. These variations in plateau stress
`follow the Clausius–Clapeyron relationship for a first-
`order transformation [16]:
`dσ
`-∆H
`Tε0
`dT
`Where dσ is the change in plateau stress, T is the test
`temperature, ∆Η is the latent heat of transformation
`(obtained from DSC measurements), and εo is the
`transformational strain. ∆Η and εo are controlled by the
`crystallography of the transformation and can be
`considered constants. The right side of the equation
`therefore defines the ‘stress rate’
`for the stress-
`induced transformations. For the present case, the
`stress rate is 6.1 MPa°C⫺1, which is within the typical
`range of 3–20 MPa°C⫺1 for Nitinol alloys [18]. The
`consequence of this relationship is that the mechanical
`properties of Ti–Ni alloys depend directly on the
`transformation temperature and test temperature.
`The ultimate
`tensile stress
`(UTS) gradually
`decreases from approximately ⫺100°C to 150°C, with
`a slight minimum at 150°C. The UTS and plateau
`
`⫽
`
`Unoading plateau
`
`Figure 4. Schematic stress–strain curve of superelastic
`Nitinol. There is a transformation from austenite to
`martensite that begins at the apparent yield stress. The
`plateau stress remains nearly constant with increasing strain
`as the amount of martensite increases. Upon unloading, the
`martensite reverts to austenite along the unloading plateau.
`The ‘permanent set’ measures any residual strain.
`
`such as guide-wires, to minimise kinking. Any
`residual strain is caused by an accumulation of plastic
`strain and is measured by the ‘permanent set’, as
`shown on the figure. Note that the stress–strain curve
`exhibits a stress hysteresis that, similar to the thermal
`hysteresis discussed above, is due to microstruc-
`tural frictional effects. The magnitude of the stress
`hysteresis plays an important role in the design of
`many Nitinol applications, such as Nitinol eyeglass
`frames. A high loading stress is required to resist
`easy bending of the frame, whereas the unloading
`stress should be low so that the temples exert a
`gentle pressure against the head. Stöckel
`[2]
`discusses other examples of this ‘biased stiffness’
`property.
`
`Effects of test temperature
`The tensile curves shown in Figure 5 illustrate that the
`mechanical behaviour of Nitinol varies greatly from
`⫺100°C to 150°C. In these tests, wires with an As of
`⫺22°C and Af of 11°C were pulled to 6% strain,
`unloaded to zero stress and were then pulled to
`failure. At the lowest test temperatures, the wires are
`martensitic and the high residual strains are fully
`recovered by heating above Af (the shape-memory
`effect). From about 0°C to 100°C the tensile curves
`exhibit superelastic ‘flags’, and we note that it
`becomes more difficult to stress-induce martensite as
`the test temperature increases. Along with the
`increase in the plateau stresses, the permanent set
`also increases with temperature. The tensile behaviour
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`Figure 5. Effect of test temperature on the mechanical behaviour of Nitinol wire. Note that there is a systematic increase in
`the loading and unloading plateau stresses with increasing test temperature. Below 0°C, the structure is martensite and,
`above 150°C, the graph shows conventional deformation of the austenite. The intermediate temperatures all show classic
`transformational superelasticity.
`
`stress converge above this temperature, which is a
`further indication that Md is near this temperature. In
`Figure 8, we see that the elongation is fairly constant
`up to about 150°C and then drops at the higher
`temperatures. The combination of the low ductility and
`high stresses above 150°C may indicate a toughness
`minimum for this material.
`
`Effects of ageing heat treatments
`Several
`investigators have shown that optimal
`superelastic performance can be achieved in Nitinol
`alloys that have a combination of cold work and
`ageing heat treatments [18, 19]. Precise control of
`these thermomechanical treatments can lead to
`reproducible mechanical properties and
`trans-
`formation temperatures. TiNi alloys with 50.8% Ni
`respond well to ageing heat treatments to ‘tune in’ the
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`Optimisation of processing and properties of medical grade Nitinol wire
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`Figure 6. Effect of test temperature on permanent set is
`shown from the data in Figure 5. The ‘superelastic window’
`extends from approximately 0°C to 60°C, where there is
`minimal residual strain after unloading from 6 % strain.
`Below 0°C, the strain can be recovered by heating above Af
`= 11°C (the shape memory effect) and at higher
`temperatures the residual strain is due to plasticity.
`
`Figure 7. Effect of test temperature on plateau and tensile
`stresses is shown from the data in Figure 5. Note that there
`is a linear relationship between plateau stress and test
`temperature from about 0°C to 100°C with a slope, or
`stress rate of 6.1 MPa/°C⫺1. The ultimate tensile stress
`shows a gradual decrease with increasing temperature with
`a minimum at the Md of 150°C.
`
`desired properties. Nishida et al. [20] established the
`effects of ageing time and temperature on the Ti-Ni
`precipitation
`reactions
`in Ti-51%Ni alloys by
`metallographic methods. They observed precipitation
`sequence of Ti11Ni14 – Ti2Ni3 – TiNi3 in the TiNi matrix
`at temperatures between 500°C and 800°C and for
`times up to 10 000 h and presented the data in a
`time–temperature–transformation (TTT) diagram. Their
`
`Figure 8. Effect of test temperature on elongation is shown
`from the data in Figure 5. The elongation does not vary
`much with temperature up to approximately 150°C. The
`drop in ductility at the high temperatures may signify a
`toughness minimum.
`
`work gave great insight into the metallurgical tool of
`controlling precipitation reactions in Nitinol alloys.
`Clearly, however, the times investigated by Nishida
`et al. [20] are significantly longer than can be tolerated
`in a production environment. Therefore, straight wires
`from the previous section were aged between 300°C
`and 550°C for 2 – 180 min, to characterise the effects
`on transformation temperature and mechanical
`properties. Figure 9 illustrates these effects on the
`Af temperature. We note that the transformation
`temperature does not change significantly at 300°C.
`Also, at 500°C, the Af increases slightly at short times,
`but does not
`increase rapidly.
`The intermediate
`temperatures, namely 350 – 450°C, have a greater
`impact on the transformation temperature. At the
`highest ageing temperature, 550°C, there is an initial
`decrease in Af and then a rapid increase. Admittedly,
`these trends of temperature and time on the Af may
`appear counter-intuitive. However, more clarity is
`gained by grouping the time–temperature conditions to
`obtain common Af temperatures. Figure 10 is such a
`TTT diagram, where each ‘c-curve’ represents the loci
`of constant Af. This figure illustrates that there is a
`maximum in the precipitation reaction at about 425°C;
`i.e., the Af increases most rapidly after heat treatments
`at 425°C. For example, the Af increases from 11°C in
`the as-straightened wire, to 30°C after ageing for only
`10 min. To reach the same 30°C Af at 500°C takes
`about 60 min and at 300°C the time exceeds 180 min.
`It is certainly beyond the scope of this article to
`review the metallurgy of precipitation reactions. How-
`ever, the shape of these curves can be understood
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`have high nucleation rates, but low diffusion rates.
`These
`two processes are optimised at
`the
`intermediate temperatures (350 – 450°C) to achieve
`rates. The Af change,
`maximum precipitation
`therefore, is due to relative Ni and Ti atom diffusion,
`where the Ni atoms congregate in the precipitates
`and the Ti atoms move to the TiNi matrix phase. As
`the matrix becomes enriched in Ti, the transformation
`temperature
`increases, as expected
`from
`the
`relationship of composition
`to
`transformation
`temperature [22]. Although the overall composition of
`the material remains Ti–50.8% Ni, localised shifts of
`composition
`can
`affect
`the
`transformation
`temperatures.
`The trends in the TTT curves indicate that a single
`precipitation
`reaction
`in
`the
`(Ti11Ni14) occurs
`temperature range 300–500°C. Between 500°C and
`550°C, however, there are ‘cusps’ in the Af curves.
`Above 500°C, the Ti11Ni14 precipitates dissolve and
`is a corresponding decrease in
`there
`the
`transformation temperature, as the Ni atoms diffuse
`back into the matrix. At 550°C the Ti2Ni3 phase
`precipitates require an even greater amount of Ni to
`diffuse away from the matrix. Precipitation of this
`phase, therefore, will again increase the Af, but at a
`different reaction rate than for Ti11Ni14. These findings
`are consistent with Miyazaki’s microstructural study of
`Ti-50.6%Ni alloys after ageing for 60 min at 400°C,
`500°C and 600°C [19]. His results demonstrate that
`the maximum density of Ti11Ni14 precipitates is
`obtained at 400°C.
`The effects of the ageing treatments on the loading
`plateau stress are shown in Figure 11. Since the wire
`was ‘strain-aged’ during the initial processing, these
`additional ageing treatments do not increase the
`loading plateau.
` Ageing temperatures
`in the
`300–500°C range systematically decrease the loading
`plateau, as we would expect with the increase in Af
`temperature. At 550°C, there is an initial decrease in
`loading plateau stress and then a more rapid
`decrease as ageing and annealing processes occur.
`The effects of ageing on the UTS are more interesting,
`as seen in Figure 12. Here the ageing treatments
`between 300°C and 450°C increase the UTS. This
`illustrates that the Ti11Ni14 precipitates are effective
`barriers to dislocation motion. Although we know that
`precipitates form during ageing at 500°C and 550°C,
`there is a dramatic decrease in UTS, especially at
`550°C. The decrease in plateau and tensile stress at
`500°C and 550°C illustrates that this is an effective
`temperature
`range
`for annealing
`(dislocation
`annihilation).
`The above ageing discussion points out that the
`transformation temperature can be readily adjusted by
`
`Figure 9. Effect of ageing temperature and time on the
`transformation temperature of Ti-50.8% Nitinol wire with a
`starting Af temperature of 11°C are shown. Note that all of
`the ageing temperatures tend to increase the transformation
`temperature, although at 550°C there is an initial decrease.
`
`Figure 10. Effect of ageing temperature and time on the
`transformation temperature of Ti-50.8% Nitinol wire with a
`starting Af temperature of 11°C are shown. The data from
`Figure 9 are re-plotted to illustrate the conventional
`time–temperature–transformation (TTT) diagram. Note that
`the maximum precipitation rate is about 400°C. Between
`500°C and 550°C the precipitates dissolve and tend to
`lower the Af. A new precipitate forms at 550°C (see text for
`more details).
`
`by briefly exploring two factors that govern diffusional
`nucleation and growth of precipitates [21]. At high
`temperatures,
`there is sufficient
`thermal energy
`to permit rapid diffusion of Ni and Ti atoms in the
`matrix. However, it becomes more difficult for the
`atoms
`to
`form a precipitate nucleus as
`the
`temperature increases. At lower temperatures,
`however, just the opposite situation occurs: here we
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`Optimisation of processing and properties of medical grade Nitinol wire
`
`diametral strain in proximal aorta at 100 mmHg
`pressure differential. Therefore, an implanted device
`could be exposed to pulsatile strains as high as 6%.
`The actual mean strain that the implant is subjected to
`depends on material and design factors. However,
`whether the stent is manufactured using woven wire,
`etched sheet, or laser-cut tubing, the strut sections
`undergo cyclic bending deformation. Although
`several papers have reported on the fatigue prop-
`erties of Ti–Ni alloy plates and bars [25] under
`tension–compression conditions, only a few recent
`papers have been published on the fatigue life of Ti–Ni
`wires [26, 27]. Furthermore, many Nitinol products are
`tested under accelerated pulsatile conditions that
`allow only periodic observations to ensure that the
`devices have not
`fractured
`(for example, see
`Reference 28). A logical approach, however may be
`to supplement these ‘submission tests’ with more
`fundamental bend–fatigue studies on wires or stent
`elements to aid in the design phase. The purpose of
`the present section, therefore, is to review the effects
`of strain amplitude, stress and test temperature on
`the fatigue life of binary Ti–Ni alloy wires.
`Miyazaki et al. tested cold-worked and aged Ti-
`50.0at%Ni
`[27] and Ti-50.9at%Ni
`[28] wires in a
`rotary-bend apparatus. The specimens were fixed in
`a bent shape with a suitable radius of curvature to
`induce a desired strain at the specimen surface at
`specified temperatures above and below Af. Figure
`fibre) strain amplitude (εa)
`13 shows the (outer
`versus number of
`rotations
`to fracture
`(Nf)
`relationship at each test temperature for the two
`alloys. The upper diagram shows three curves for
`the Ti-50.9at%Ni alloy and the data for
`the Ti-
`50.0at%Ni alloy is in the lower diagram. Both
`alloys show a general trend of increasing fatigue life
`with decreasing test temperature in the highest and
`intermediate strain-amplitude regions. In the higher Ti
`alloy the fatigue-endurance limit increases with
`decreasing temperature below Af. The fatigue limit is
`insensitive to temperature for the Ti-50.0at%Ni alloy
`above Af and for all conditions for the Ti-50.9at%Ni.
`Miyazaki et al. [27] carefully studied the details of
`tensile stress–strain curves, to gain insight into the
`fundamental mechanisms that influence fatigue
`behaviour and give rise to the differences observed in
`Figure 13. They measured the proportional stress
`limit (σpr) and the critical stress to induce the
`martensitic transformation (σM) and corresponding
`strains, as schematically illustrated in Figure 14.
`Below the proportional limit strain (εpr) there is pure
`elastic deformation, whereas between εpr and the
`elastic limit strain (εe) there is an elastic deformation,
`including microscopic local twinning or microscopic
`
`Figure11.Theeffectoftimeandtemperatureontheloading
`plateauofagedTi-50.8%Nitinolalloy.Thesestressdata
`correspond to theAftemperaturesshowninFigures9and10.
`Thereisasystematicdecreaseinplateaustresswithincreasing
`temperature.Thereisamoredramaticeffectat550°C.
`
`Figure 12. The effect of time and temperature on the
`ultimate tensile stress. Ageing temperatures between 350°C
`and 450°C tend to increase the tensile strength due to
`precipitation hardening. At 500°C and 550°C, the annealing
`effects dominate and lower the strength.
`
`selecting an appropriate time and temperature. Higher
`Af temperatures are achieved by ageing in the 300 –
`500°C range. Additionally, the Af can be lowered by
`short ageing times between 500 and 550°C.
`
`Fatigue properties
`Fatigue life is a major concern for biomedical implant
`applications. For example, the US FDA requires proof
`of fatigue resistance of 10 years (400 million cycles) in
`simulated body environment for intravascular stents
`[23]. Pedersen et al.
`[24] measured 6% average
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`A.R. Pelton et al.
`
`growth rates were much slower in fully martensitic
`Ti–Ni alloys than in alloys that undergo a stress-
`induced transformation.
`Since Nitinol alloys perform better under strain
`control than under load control, most fatigue studies
`plot fatigue life as a function of stra