5-?--qå
AN INVESTIGATION INTO PHYSICAL AND MICROSTRUCTURAL PROPERTIES OF ORTHODONTIC WIRES AND CHANGES INCIDENT TO BENDING.
A research report submitted in partial fulfilment of the requirements for the degree of Master of Dental Surgery
by
S.L. WHITTLË B.D.S. (Adel)
Department of Dentistry' FacultY of Dentistry' The UniversitY of Adelaide, South Australia.
December, 1995 tl
TABLE OF CONTENTS
Page No.
LIST OF FIGURES tv LIST OF TABLES ix LIST OF ABBREVIATIONS xii SUMMARY xiii SIGNED STATEMENT xvi ACKNOWLEDGEMENTS xvii
1 CHAPTER 1 INTRODUCTION AND AIMS
CHAPTER 2 LITERATURE REVIEW
HISTONICRL ASPECTS OF ORTHODONTIC WIRES 4 PHYS]CAL PROPERIES OF WINES 5 Stiffness 7 Strength 10 Springback 11 FormabilitY 12 JoinabilitY 13 BiocomPatibilitY and StabilitY 13 Friction 14 Sretnless Sreel WIRES 15 SreruoRnos FoR OnrHoooNrc WIRES 26 BENDS IN ORTHODONTIC WINES 32
CHAPTER 3 MATERIALS AND METHODS
PRepRR¡TION OF WIRES FOR SEM Materials 47 Methods 52 Teuslu Teslruc or WtlcocKrM WIRES Materials 67 Methods 69 ARCH.FORM ANCHORAGE BEND ANALYSIS Materials 72 Methods 82
CHAPTER 4 RESULTS AND DISCUSSION
THE SEM EXAMINATION OF WINES Wire microstructure 85 Bending effects from different types of pliers 94 Defects and lnclusions 109 Fractures 118 131 TENSILE TESTING OF WILCOCKTM WIRES ARCH.FORIU AUCHONAGE BEND ANALYSIS 142
CHAPTER 5 CONCLUSIONS
161 THE MICNOSTRUCTURE OT WINES Teruslrc Teslruc or WtlcocKrM WIREs 163 165 ARCH.FORM ANCHORAGE BEND AruRIYSIS
CHAPTER 6 FUTURE RESEARCH 167 THE MTCROSTRUCTURE OF WIRES 168 TENSILE TESTINC OT WIICOCKTM WIRES ARCH.FORM ANCHORAGE BEND ANALYSIS 169
170 CHAPTER 7 BIBLIOGRAPHY
185 APPENDIX 1 PHOTOELASTIC MODELING iv
LIST OF FIGURES
FIGURE SUBJECT Page No
2 - 1 'Face centred cubic' (fcc) lattice. 16
2 -2 'Body centred cubic' (bcc) lattice. 16
2-g Typical wire drawing die, (Dieter, 196f ). 20
2 - 4 lllustration of terms used in the theory of plastic bending. 35
2-5 Graphical representation of variation in intrusive force as a function of degree of anchorage bend as measured by several authors. 42
2-6 Graphical representation of variation in intrusive force as a function of distance of a 45 degree anchorage bend from the buccal molar tube as measured by several authors. 44
3- 1 Wire drawing by A.J. Wilcock Scientific and Engineering Equipment, (with kind permission from Mr. A.J. Wilcock). 48
3 - 2 Anchorage bend forming guide. 53
3 - 3 Detail of beaks of stainless steel tipped light wire pliers. 54
3-4 Detail of beaks of stainless steel tipped Tweed torquing pliers. 55
3-5 Detail of beaks of tungsten carbide tipped light wire pliers. 56 v
F¡GURE SUBJECT Page No
3-6 Detail of beaks of tungsten carbide tipped high tensile wire pliers. 57
g-7 Abramin (Struers) grinding and polishing machine. 59
3 - I Surface of 0.016 inch, (0.4064 mm) 'premium plus' Wilcock wire, batch 'TT-TAT', following grinding, polishing, and etching. 64
3 - 9 Philips XL20 Scanning Electron Microscope. 65
3 - 1O Specimen stage of Philips XL20 Scanning Electron Microscope. 66
3 - 11 Hounsfield Vee Wire GriPs. 70
3 - 12 lnstron Universal Testing Machine. 71
3 - 13 The representative arch form. 75
3 - 14 Arch-wire designs. 77
3 - 15 CAD drawing of the arch-wire testing apparatus. 78
3 - 16 View of apparatus with an activated arch-wire 79
3 - 17 Arch-wire anchorage analysis apparatus 81
4 - 1 X-ray micro-analysis of the main bulk of wire in batch .TT-TAT" 88
4-2 Schaeffler diagram showing effect of alloying elements on the basic structure of chromium - nickel steels, (from Honeycombe, 1981) 90 vl
Page No FIGURE SUBJECT
4-3 Surface irregularities in the vicinity ol a 40 degree bend in 'as-received' TT-TAT, 0.016 inch, (0.4064 mm) 'premium plus'. 95
4-4 Specimen showing increased pitting along a longitudinal 'core'. 97
98 4-5 Specimen displaying reduced etching effect
4-6 Specimen displaying increased etch¡ng effect' 99
4 - 7 Stainless steel light wire plier effect on concave surface of a 50 degree bend. 101
4-8 Example of plier edge defect in 30 degree specimen 103 bent with Tweed torquing Pliers.
4-9 Example of plier edge defect in 40 degree specimen 105 bent with tungsten carbide tipped, light wire pliers'
4-loPlieredgedefectasviewedfromthesideofan incompletelypolishedspecimenpreparedwithtungsten 106 carbide tiPPed light wire Pliers.
4 - 11 Typical example of specimen prepared with 'dolphin' 108 tungsten carbide tiPPed Pliers.
4 - 12 Longitudinal surface markings from the wire drawing processof0.Ol6inch(0.4064mm)Wilcock.premium 110 Plus', batch TT-TAT
4 - 19 Photomicrograph of Al - Ti - o inclusion identified in batch'TT-TAT', 112 vii
Page No FIGURE SUBJECT
4-14 X-ray micro-analysis of an inclusion found in batch .TT.TAT' '114
115 4 - 15 Example of polishing particles.
4 - 16 Deliberate fracture of 0.016 inch (0.4064 mm) wilcock 121 'Premium Plus', batch'TT-TAT'.
4 - 17 Detail of deliberate fracture of 0.016 inch (0.4064 mm) Wilcock'premium plus', batch'TT-TAT" '122
4 - 18 Further detail of deliberate fracture of 0.016 inch (0.4064 mm) Wilcock'premium plus', batch 'TT-TAT'' 123
4 - 19 Longitudinal failure ('F') seen in specimen bent at 50 degrees with 'Dentronix 105' pliers' 125
4 - 20 Detail of longitudinal failure seen in specimen bent at 50 degrees with 'Dentronix 105' pliers' 126
4 - 21 Radiating lines ('R') of deformation that may represent a pre-fai lure condition. 128
4 - 22 Further detail of a radiating line of deformation that may represent a pre-failure condition. 129
4-23 Example of a transverse failure of high tensile wire possibly resulting from transverse strain inherent in the wire, (Lee, 1992), (with kind permission from the Australian Orthodontic Journal) 130
4 - 24 Analysis of test arch-wires with 15 degree anchorage 145 bends. vill
FIGURE SUBJECT Page No
4 - 25 Analysis of test arch-wires with 30 degree anchorage bends 146
4 - 26 Analysis of test arch-wires with 45 degree anchorage bends. 147
4 - 27 Analysis of test arch-wires with 60 degree anchorage bends. 148
4 - 28 Analysis of 'plain' arch-wires with varying degree anchorage bends. 153
4 - 29 Analysis of inter-maxillary circle' arch-wires with varying degree anchorage bends. 154
4 - 30 Analysis of 'boot hook' arch-wires with varying degree anchorage bends. 155
4-31 Analysis of 'loop hook' arch-wires with varying degree anchorage bends. 156
4 - 92 Analysis of 'molar stopped' arch-wires with varying degree anchorage bends. 157 tx
LIST OF TABLES
No TABLE SUBJECT Page
2-1 Classification of wires (by strength) as required by the Standards Association of Australia, AS 1964 - 1977 lor orthodontic wires, (resilient),(excluding precious metal wires). 27
2 - 2 Yield strength requirements as required by the American Dental Association specification Number 32 for orthodontic wires not containing precious alloys. 27
2-3 Anterior intrusive force from 45 degree anchorage bends placed in 0.016 inch (0.4064 mm) Wilcock wires as measured by four separate authors. 40
2- 4 Anterior intrusive force from anchorage bends of varying degree placed in 0.016 inch (0.4064 mm) 'special plus' Wilcock wire as measured by several authors' 41
2-5 Anterior intrusive force generated by 0.016 inch (0.4064 mm) Wilcock'special plus' arch-wires with 45 degree anchorage bends as a function of distance of anchorage bend from the buccal tube as measured by several authors. 43
3-1 Diameter and tensile properties of batch 'TT'TAT' of Wilcock 0.016 inch, (0.4064 mm) 'premium plus" 49
3-2 Normal elemental composition for AlSl Type 302 stainless steel, (ASTM ,1990), compared to 'Australian' wire as determined by Newman, (1963)' 51 X
No TABLE SUBJECT Page
3-3 Details of specimens prepared with various plier types 52 and magnitude of bend Placed.
3-4 Grinding and polishing protocol for preparation of wire specimens. 58
3-5 Comparison of orthodontic wire types, as classified by the Australian Standard, AS 1964 '1977, and A'J' Wilcock Scientific and Engineering Equipment (AJW)' 67
3-6 Tensile strength of Wilcock wires as measured by A'J' Wilcock Scientific and Engineering Equipment (AJW) and the Australian Dental Standards Laboratory (ADSL). 68
g - 7 Maxillary cast measurements to determine arch-form dimensions, in millimetres, compared with the results of Borghesi, (1973) and Rohan, (1982)' 73
3 - 8 Degrees of anchorage bend investigated by previous authors. 83
4-1 X-ray micro-analysis of electron 'K' shell energy to determine relative concentration of elements in A'J Wilcock batch TT-TAT 0.016 inch (0.4064 mm) 'premium plus' 87
4-2 Compositional analysis of 'Australian' wire by Newman, (1963). 92
4-3 Results of each valid tensile test, (T1 to T6), and the resultant mean tensile load for failure for each batch of specimens. 132 xi
TABLE SUBJECT Page No
4-4 Tensile strength of Wilcock wires as measured by A.J. Wilcock Scientific and Engineering Equipment, (AJW), the Australian Dental Standards Laboratory, (ADSL), and by experimentation, (SLW). 132
4-5 Comparison of orthodontic wire classifications, as classified by A.J. Wilcock Scientific and Engineering Equipment, (AJW), Australian Dental Standards Laboratory, (ADSL), and during experimentation, (sLW). 133
4-6 Comparison of wire tensile classification as calculated with actual versus nominal diameters. 134
4-7 Percentage differences in tensile strength as tested by the Australian Dental Standards Laboratory, (ADSL), and experimentation, (SLW) when compared to results of A.J. Wilcock Scientific and Engineering Equipment, (AJW). 138
A-1 Elastic (Young's) modulus for components of the stomatog nathic system. 186 xil
LIST OF ABBREVIATIONS
ABBREVIATION DEFINITION
ADSL Australian Dental Standards Laboratory
AISI American lron and Steel lnstitute
AJW A.J. Wilcock Scientific and Engineering Equipment
ASTM American Society for Testing Materials
bcc Body centred cubic CAD Computer Aided Design CEMMSA Centre for Electron Microscopy and Microstructural Analysis, The University of Adelaide.
E Young's modulus of elasticitY
fcc Face centred cubic
kg Kilograms
KV Kilovolts pA Microamperes pm Micrometres
mm Millimetres
MPa Megapascals
N Newtons
nA Nanoamperes
Pa Pascals PDL Periodontal ligament SEM Scanning Electron MicroscoPe x ¡¡¡
SUMMARY
Anecdotal accounts of an unacceptable breakage rate of high tensile sta¡nless steel orthodontic wires have been received by their manufacturer, A.J. Wilcock of Whitttesea, Victoria. lt would seem that these allegations are directly related to probtems in the successful clinical manipulation of the wire with orthodontic pliers. lt was decided to investigate some of the bending characteristics of the high tensile 0.016 inch (0.4064 millimetre) 'premium ptus' grade Wilcockl wire to identify any cause lor a higher than expected breakage rate.
Those physical properties considered ¡mportant in the behaviour of orthodontic wires are: stiffness, strength, springback, formability, joinability, biocompatibility and stab¡lity, and friction. These physical properties have been thoroughly investigated, and form the basis for lndustrial Standards in a number of countries. However, an examination of the literature concerning stainless steel orthodontic wire reveals disagreement as to which tests best describe the requirements of such fine wires for clinical orthodontic purposes.
Tensile testing of aged wire specimens has indicated that strength may change over time. Results from this investigation found that ultimate tensile strength for aged specimens was up to ten per cent greater than the ult¡mate tensile strength determined from the same batches of wire by the Australian Dental Standards Laboratory in 1985. Thus the ageing of wire may increase
1 n.¡. Wlcock Scientific and Engineering Equipment, Whittlesea, Victoria, Australia. xlv its strength, although, differences in the subtleties of testing technique cannot be eliminated as a cause of these differences.
The structure of fine, high tensile ofthodontic wires is considered to contribute significantly to their behaviour but has received very little attent¡on in the availabte literature. This is particularly evident in the inability to calculate Young's modulus accurately (ie: a numerical expression of a materials' retative rigidity). 'As-received' wire was used to produce specimens that simulated clinical manipulation by bending. These were then prepared by longitudinal grinding and polishing and subsequent etching. The internal diameter of these specimens was viewed by Scanning Electron Microscopy to investigate the microstructure in and around the bent region.
Different pliers were used to place bends of varying degrees in a series of wire specimens, and the microstructural features were compared across the diameter of straight and bent portions of the wire. When compared with stainless steel tipped pliers, unmodified tungsten carbide tipped pliers showed a more severe deformation on the concave, internal region of the bend. Modified tungsten carbide tipped pliers however, showed little if any damage to this same region of the wire
No breakages were found during specimen preparation when using a standard clinical bending technique. Therefore, in order to study fracture behaviour, forces inconsistent with those used during normal clinical manipulation were deliberately used to fracture some wire specimens. This was best achieved by bending the wire very quickly with a snapping motion. Scanning Electron Microscopy suggests the wire has a lamellar structure, and that specimen failure is consistent with the 'peeling' seen clinically when faulty manipulation causes unwanted fractures. 'Fast'fractures of this type seem to XV or¡ginate on the convex, external surface of the bend. A comparison was made between this type of fracture pattern and those seen in some specimens collected in the clinical situation, where the initial failure point was also found on the external, convex surface of the bend.
It is suggested that to avoid fractures in high tensile stainless steel wires, two factors in manipulation are critical. Firstly, bends should be made slowly with the fingers. Secondly, unmodified tungsten carbide tipped pliers should be used with cautiorì, or, the edges of the square beaks modified by smoothing with a cratex tyPe wheel.
Force levels at different points along an arch-wire were measured as a function of the various types of 'canine hooks' employed in the Begg technique, and, according to the amount of anchorage bend placed in the wire. Results suggest that different types of 'canine hooks' have negligible effect on anterior force values, but, consistent with beam theory, forces exhibited in the premolar region are about ten times greater than in the incisal region. These results suggest that caution should be exercised when part attaching premolars to an arch-wire designed for anterior intrusion as a of the Begg technique. This same caution should be exercised if molar tipping is undesirable. xv¡
STGNED STATEMENT
This report contains no material that has been accepted for the award of any other degree or diploma in any other university or tertiary institution. To the best of my knowledge and belief, it contains no material previously published or written by another person, except where due reference is made in the text.
The views and opinions expressed herein do not necessarily state or reflect those of the Government of the Commonwealth of Australia, The Department of Defence, or, the Royal Australian Navy.
I give consent to this copy of my report, when deposited in the University of Adelaide Library, being available for loan and photocopying in accordance with existent copyright conditions.
Samuel L. Whittle.
tr;$L"a day of 2"..*U 1 995 xvil
ACKNOWLEDG EM E NTS
I wish to express my sincere gratitude to the following:
Dr John V. Bee, Senior Lecturer; Department of Mechanical Engineering, The University of Adelaide, who as a supervisor, readily gave of his expertise, practical advice, and good humour.
Dr Wayne J. Sampson, Senior Lecturer in Orthodontics, Department of Dentistry, The University of Adelaide, who, aS a supervisor, inspired in me, a fascination for investigation and research.
Mr lan Brown, Department of Mechanical Engineering, The University of Adelaide, for valuable information and his assistance in testing techniques.
Mr John Terlett, Centre for Electron Microscopy and Microstructural Analysis, The Univerçity of Adelaide, for his technical assistance in using the Scanning Electron Microscope.
V¡kki Hargreaves, Margaret Leppard, Peter Dent, and Raoul Pietrobon, Department of Dentistry, The University of Adelaide, for their assistance in experimental and photographic techniques.
Arthur Wilcock and Ella White of A.J. Wilcock Scientific and Engineering Equipment, Whittlesea, Victoria, for their ready provision of test material, valuable advice and hospitality.
Garth L. Brice, Orthodontist, of Chapman, Australian Capital Territory, for his role as mentor and true friend during my career in the Royal Australian Navy.
My son Hamish, whose love and ready smile made it all worthwhile.
To my dear wife Heather, without whose unending patience, support and devotion, nothing of this venture would have been possible. CHAPTER 1
INTRODUCTION AND AIMS
The goal of modern orthodontics can be summed up as the creation of the facial best possible occlusal relationships within the framework of acceptable æsthetics and stability of the occlusal framework", (Proffit, 1993)' one generally avenue to achieve this aim is by the use of fixed appliances, to comprising some form of attachment to the teeth, and some mechanism allow mechanical therapeutic interaction of the teeth. All fixed orthodontic principles of appliances rely on the use of wires to move teeth and depend on force production derived from such simple machines as the lever, the spring' from the inclined plane, the screw, the wheel and axle, and the forces derived the elasticity of rubber, and the resitiency of wire, (McCoy & Shepard' 1e56).
The principles of tooth movement dictate that a mechanical system to alter the point arrangement of the teeth must be capable of delivering both single (tipping) and double point (couples) forces, (Gianelly & Goldman, 1971)' "Little emphasis has been placed on the physical manipulation of (Begg) delivered to archwires whereby varying magnitudes of intraoral forces can be the periodontal structures according to the manner in which the appliance is paucity of constructed and then activated in the mouth", (Sims, 1977)' This relation information in the literature has been addressed of late, particularly in to the use of new generation elastic and super-elastic wires' However' a search through the literature reveals that there has not been an equivalent output in relation to conventional stainless steel orthodontic wires. 2
Perhaps the singutarly most important component of the Begg orthodontic technique is the anchorage bends placed in the archwire mesial to the molar buccal tubes. By the action of the anchorage bend, the application of is made differential forces to the teeth is possible. lntrusion of anterior teeth possible by anchorage bends, as is the resistance to the anterior component
of force exerted by intermaxillary elastics acting on molars.
ln the production of wires for orthodontic purposes, the variation in concentration of alloying elements and the manipulation of manufacturing technique allows for fine control over the microstructure of the wire' thus Consequently, it is possible to modify the mechanical properties, and quality clinical behaviour, of the wire. lt has been suggested that the of the orthodontic wires is not always within clinically acceptable limits, despite practical difficulties in manufacture, (Williams & Fraunhofer, 1971)' The its clinical performance of an orthodontic wire is strongly dependant upon However' manufacturing process, clinical manipulation, and heat treatment' The ,apparently standard techniques produce erratic clinical performances"' the reason behind this discrepancy is that the critical relationship between composition, structure, and the attendant physical properties is not currently well understood. This is understandable given the relationship is not only complex, but difficult to investigate, (¡b¡d).
The broad aim of this investigation was to investigate orthodontic stainless steel wire subsequent to bending. Given that the anchorage bend is some of fundamental in Begg orthodontic treatment, it was decided to explore and the properties and behaviour of wires in the vicinity of anchorage bends' where possible relate this to clinical maniputation and use of the wire' 3
The specific aims were as follows:
vary¡ng a To investigate the microstructure of wires that had been bent to degrees;
a To investigate the effect on microstructure that bending with different
commonly used clinical Pliers;
a To consider the relevance of applicable industrial standards to the clinical manipulation of orthodontic stainless steel wires; the a To investigate the significance of extended storage of wires, and resultant effects, if any, on physical properties, and on tensile strength in particular;
a To investigate force levels exefted by typical 'Begg stage 1' arch-wires'
Where possible, the investigations were kept clinically relevant
ì
I 4
CHAPTER 2
LITERATURE REVIEW
The earliest orthodontic appliances of Fauchard (1746) were fabricated with short strips of gold or silver (bandelettes) tied to teeth with lengths of waxed thread. The use of precious alloys as a basis of orthodontic appliances continued to prevail through the era of Edward Hartley Angle, who's'E' arch in with 1903 was constructed Of heavy, round, German silver wire, and continued the introduction of the edgewise appliance in 1928. Calvin Suveril Case was one of the first people to use a light round wire for looth movement, although the properties of the precious alloys in use at the time made such treatment very difficult, (Dixon, 19721-
its The development of stainless steel by Brearly of Sheffield in 1913, allowed introduction into orthodontics in France by the Belgian, De Coster in 1932' (¡b¡d). lt was subsequently championed by a graduate of Angle's school' sheldon Friel, (Friel, 1933). He described a stainless steel appliance and suggested stainless steel to be superior to precious alloys due to reduced cost, greater tensile strength, greater toughness, less likelihood to tarnish in the mouth, and less likely to retain debris due to 'miniaturisation'. He stated harden the need to use austenitic stainless steel due to the ability to cold work to the material. The disadvantage of stainless steel was cited as being difficult manipulate as compared to precious alloys. However, Friel described a method of welding which is similar to that in use today, (¡b¡d). 5
PHYSICAL PROPERTIES OF WIRES
The ideal orthodontic wire should move teeth by light continuous forces. lt should have the properties of low stiffness, high strength, high range of etasticity, high formability, be weldable or sotder-able, (Proffit, 1993)' be resistant to corrosion in the mouth, be biocompatible and environmentally stable, have low surface friction (Kapita & Sachdeva, 1989, Singh' 1993), and be of reasonable cost, (Proffit, 1993). An appreciation of a wire'S properties can be easily assessed by the feel of the wire in one's hands. lt should flex easily without "taking a set", should be formable into an arch shape with relative ease, and should be resilient, or have good springback, (Wilcock, 1989). ldeally, thé light, continuous force should be exerted over large displacements without undergoing permanent deformation, (Wilcock, 1988). There is no universal agreement to which properties or set of tests are the most important in characterising orthodontic wires, however, it is generally agreed that tests that describe the elastic behaviour of wires are the most relevant. Elasticity however can be described by an array of terms, and measured by a wide range of tests, (Twetftree, 1974)' The majority of tests allow for a comparison of different wires, but may not be clinically subjective, (Kapila & Sachdeva' 1989). Those properties that affect the clinical selection of wires should be understood.
All orthodontic tooth movement is accomplished by the application of force to the teeth by some mechanism that has the capacity to store energy and deliver such energy to the teeth in a predetermined way. Force levels considered optimal for tooth movement are acknowledged as being in the range of 0.76 - 1.26 N, (Quinn & YoShikawa, 1985), or, when based on root surface area, between 0.015 and 0.026 N / mm2, (Lee, 1965). Force levels of 0.020 - 0.026 N / mm2 were first suggested by Schwarz in 1932 as 6 being the threshold level for the initiation of root resorption, (Brezniak & wasserstein, 1993), however this figure has since been upgraded to 0.08 N / mm2 by Miura, (1973). Those who use the bioprogressive technique suggest forces as light as 0.01 N / mm2 are suitable for tooth movement, (Ricketts et al, 1979). lt has been suggested that force per unit areaol tooth root as it bears on the periodontal ligament, and thus bone, has greater significance than force levels exerted on the tooth itself by an appliance, (Storey & Smith, 1952).
Basic elastic properties that allow this energy to be delivered are stiffness, or resistance to deformation; strength, or the ability to sustain a load; and working range, or springback. These characteristics are not necessarily inter-propottional, but are dependent on each type of wire and must therefore be considered individually when selecting a wire for a particular application. Stiffness is an indication of the rate of force delivery, whereas strength and
range arø a measure of the maximum capacity of the wire, (Thurow, 1972)' The means by which one applies a wire to an orthodontic situation, or the application of mechanical design, will determine what effect stiffness, strength, and range will have on the teeth, (Adams et al, 1987). Other related mechanical characteristics of orthodontic arch-wires are well documented in standard engineering texts, the scientific literature, and thesis presentations elsewhere, (Borghesi, 1973; Twelftree, 1974; Wiltiams, 1974i Rohan, 1982; Gullotta, 1985; Squires, 1991 ; singh' 1993; Ailen, 1gg4). The multitude of other inter-related properties are often explained by a mixture of terms and will not be repeated or elucidated here in order to avoid confusion. 7
STt F FN ESS
Stiffness is the deflection of a wire under a load and is proport¡onal to the modulus of elasticity ('E'), or Young's modulus, and connotes rigidity' (phillips, 1g73). The modulus of elasticity has traditionally been regarded as an invariant property for a particular material, (Kusy & Dilley, 1984)' calculation is based on beam theory methodology, using the following equation:
P.L3 D= 3.E.1
where: D= deflection at the free end of a cantilever
t-Þ: load at the free end of the cantilever
t- L beam length
E L apparent flexure modulus of elasticity l= moment of inertia.
,l' is a function of the beam geometry and relevant equations relative to a particular situation are usually obtained from standard engineering texts, (Dieter, 1961). For example, round wires of a given diameter can have a value for'l'derived from the following equation:
where: | : moment of inertia d = diameter of the round wire, (Kusy & Greenberg, 1981) I ln a recent paper, (Yoshikawa et À1, 1981),'E'values for 0'016 inch, (0.4064 mm) round stainless steel wires were calculated as 15.5 x 104 MPa. This value is about 20 per cent below published data that suggests values of 19.3 to 20.0 x104 MPa, (¡b¡d). Values for'E'as high as24'8 x 104 MPa have been reported, (Goldberg, Burstone & Koenig, 1983).
The effects of cold drawing may account for this apparent reduction in 'E' vatues, (Goldberg et at, 1977; Kusy & Ditley, 1984), however the opposite effect was noted for cold drawn B-titanium wires, (Shastry & Goldberg, 1983). The results of Goldberg et al, (19771 were subjected to linear regression analysis and validated values for'E'. However, rather than the usual stress-strain derivation of 'E', beam theory is used because of the ¡nherent anisotropic tendency of heavily drawn orthodontic wire, (ie, the grains of the alloy assume directional properties when heavily cold worked), and therefore, there is questionable validity for flexural data derived from tensile data. Different values of 'E' for a wire will affect design criteria for springs, (Goldberg, Burstone & Koenig, 1983). There are difficulties in accurately measuring the modulus of elasticity for thin wires, (Brantley et al, 1978; Goldberg, Burstone & Koenig, 1983; Allen, 1994). This difficulty may be attributable to the reliance of beam theory In calculation, or the failure of most materials to follow a purely linear stress-strain relationship during elastic deformation, (Twetftree, 19741. Alternatively, it has been suggested that differences in published values for tensile data are likely to be due to technique difficulties. The smallest slippage of specimens in a tensiometer during measurement can give false readings, and thus inaccuracies in calculating 'E' values, (Delgado & Anderson' 1963)'
Kapita & sachdeva, (1989), suggest that orthodontic wires should have low stiffness, or load deflection rates, to allow for: I
a "The ability to apply lower forces,
a a more constant force over time as the appliance experiences deactivation,
a greater ease and accuracy in applying a given force"'
Stiffness is important because, "all other things being equal, it describes the 'force to activation' ratio of an appliance", (Kusy et al, 1988). lf a lighter force is required of an appliance, (ie, to reduce the force per unit activation), then this can be achieved bY:
a changing only the alloy (and thus 'E'),
a setecting a different geometric size or shape (ie, altered moment of
inertia - 'l'),
a varying both ('E'and'l'), (¡b¡d).
Selection of a wire based on its stiffness, or, load-deflection characteristics, determines the type of force available to move a tooth, (Burstone' 1981)' Borghesi, (1973) suggests that many quantitative studies have been done in relation to load-deflection characteristics of orthodontic appliances, but, that many of the studies cannot be inter-related because of individual differences in materials and methods. Such a poor correlation leads to difficulty in drawing valid conclusions. Theoretically however, for a given clinical situation, any load should provide a low, continuous, physiologic force, (Johnson & Lee, 1989). 10
STR ENGTH
Strength is related to the maximum stress required to fracture a wire and is proportional to the wire's resiliency. ln orthodontic wire terms, the important factor is the point at which a wire ceases to behave elastically and begins to undergo plastiq deformation. This is defined as the yield strength, where, for a given amount of stress, there is a 0.1 per cent or 0.2 per cent permanent strain, (Greener et ât, 1972). The yield strength of a wire can be increased by:
a work hardening,
a dislocation locking,
a solid solution strengthening,
a grain refinement and orientation.
It is possible to achieve yield strengths in excess of 3.1 x 103 MPaduring the manufacture of stainless steel wires. These wires will have a high density of dislocations, will have anisotropic tendencies due to high drawing tensions, and if handled incorrectly will have a high incidence of fracture, (Wilcock' l g3g). The obvious advantage in such high yield strength wires, is the increased resiliency. This allows for stainless steel wires with a lower load-
deflection ratio, indicative of a greater constancy of force levels over distance. A constancy of force levels over distance is cited as an advantage when moving teeth, since fewer appliance adjustments are required, (Burstone et À1,1961).
Stainless steel wires are cold worked to provide high yield strength. This, coupled w¡h a high modulus of elasticity, allow for wire activat¡on over long distances. Commercially available stainless steel wires show a range of 11 values for both the modulus of elasticity and yield strengths, (Kusy et al, l ggg). Commercially processed and tested materials cannot be assumed to have been under ideal conditions when determining the definitions of 'elastic limit' and'propottional limit'.
Hooke's law is given by: 6=E.€
where: o: stress (tensile or compressive),
l- L Young's modulus of elasticitY, strain,
Deviations from Hooke's law, and slight deformations will result in varying results for yield strength, (Phillips, 1973). For most metals Hooke's law is only strictly valid for a narrow range of stress and strain values, (Dieter' 1961). Additionally, such a simplistic analysis of elasto-plastic deformation, such as occurs in wire bending, is of limited applicability, (Waters, 1972)'
SPRINGBACK
Springback, (also known as range, working range, maximum flexibility' range of activation, range of deflection, or maximum elastic deflection) is related to the ratio of yield strength to the modulus of elasticity of the material, (Kapila & Sachdeva, 1989), and is an indication of the amount of bend¡ng that can be tolerated in a wire, (Thurow, 19721. The limit of springback is defined as the proportional limit, where permanent strain results from applied stress, (phillips, lg73). Springback therefore is related to resiliency and a function of the wire's stored energy when activated. 12
Mechanical theory would suggest that springback lor a solid wire should decrease as its size increases. This was at variance with the results of lngram et al, (1986), who found that springback properties tended to be similar for a series of diameters of given wire types, notably multi-stranded staintess steel and nickel-titanium wires. Given the variety of alloys in which wires are availabte, the use of a 'stiffness number' has been suggested, (Burstone, 1981). This allows a comparison of the springback of different wires of different compositions. Different compositions between wires of a similar diameter, can have a marked effect on the¡r springback quality, (Barrowes, 1982).
FORMABILITY
Formability is the ability to form the wire into various shapes, as required to effect tooth movement, without fracture of the wire. This might take the form of bends, loopS, coils, etcetera. The use of 'V' bends as a part of force systems is common, (Bequain, 1992; Burstone & Koenig, 1976 & 1998; Ronay et at, 1989). Loops are also used in many appliances, (Lane & Nikolai, 1980), often to increase the flexibility of an initial aligning arch, (Waters, 1976 a, 1976 b). The ability to form orthodontic wires will in part depend on the amount of work hardening of the wire during manufacture. Excessive work hardening, and thus a reduction in ductility, or capacity for plastic deformation, will make the wire more liable to fail by fracture during clinical manipulation. ln theoretical terms, high formability requires an ability to effect large amounts of plastic strain within the ultimate strength of the material. 13
JOINABILITY
Joinability is the ability to weld or solder a wire to a desirable configuration without a concomitant reduction in desirable physical properties. This includes the addition of auxiliaries that may be required to modify or augment treatment, (Kapila & Sachdeva, 1989; Proffit, 1993). The ability to wetd staintess steel wires by electrical resistance methods was introduced by Friet, (1g33). A short time later, he assisted in the development and use of fluoride flux as a method of soldering stainless steel.
The temperature of the applied heat is the important variable in soldering, rather than the time of heating, (W¡lkinson, 1963). An excessive temperature in soldering or welding will cause annealing of the wire, and a subsequent loss of those properties desirable for an orthodontic wire, ie stiffness, strength, and springback. The 'maximum current for minimum time' rule applies to welding, (Friel, 1933), since this allows for a concentration of high heat at the junction of the components to be fused, and limiting any annealing of the metal.
BIOCOMPATIBILITY AND STABILITY
This includes the resistance of wires to corrosion in the oral environment and the absence of allergies or tissue intolerance. Stability reflects the need for wires to maintain their properties for extended periods after manufacture and to act in a predictable fashion over time, (Kapila & Sachdeva, 1989), and incorporates reducing potential stress relaxation to a minimum, (Twelftree et al, 1g771. The precise control required of an appliance by an orthodontist dictates that the elastic properties and dimensional stability of formed arch- 14 wires should be predictable, (Howe et al, 1963). coil springs exhibit a decay rate in force levels after activation, (Angolkar et al' 1992)'
found to Annealing in the range of 315 - 540 degrees Celsius has been residual provide satisfactory 'stress relief'. This process is claimed to reduce following stress in wires by ressening the non-uniform prastic strain that exists wire drawing, (ibid).
F RICTIO N
Friction is the result of forces when there is relative motion between two surfaces are bodies in contact, or the potential for it, and when the contacting and not perfectly smooth, (Frank & Nikolai, 1980). Thus there is a static kinetic component of the frictional force. Activation of an orthodontic move in appliance must overcome these frictional forces if teeth are to
response to the aPPlied force.
on The relative roughness of two surfaces in contact is essentially dependent by the the absolute roughness of each individual surface. This is determined etcetera)' materiat itself, the manufacturing process, (polishing, heat treatment' or creep' and to shelf life deterioration effects, such as corrosion, relaxation,
of teeth in Free tipping of teeth in light wire techniques and bodily retract¡on Excessive edgewise techniques involves a motion of bracket over wire' in a need for amounts of friction between these two components will result ideally higher force levels and greater demands on anchorage' Wires should not bind with brackets, (Kapila & Sachdeva, 1989)' 15
STAINLESS STEEL WIRES
per To produce a stainless steel, the chromium content must exceed 11 cent' lf the chromium content is greater than 11 per cent, then small amounts of other carbon are needed ¡f ¡t is to remain 'stainless'. lt may contain elements of than iron, carbon, and chromium. These alloys resist tarnishing because the presence of chromium that forms a very thin impervious oxide layer on the surface, (Philtips, 1973). Stainless steel alloys include a vast range of lnstitute compositions and physical properties. The American lron and steel (Alsl) have instituted a numbering system for alloys and stainless steel based on the elements contained in the steel, such as carbon, nitrogen, silicon, phosphorus, titanium, niobium, manganese, cobalt, nickel, or molybdenum' steel, Each of these elements has the capacity to alter the properties of the such as strength, weld-ability, or the effects of heat treatment, depending on (Peckner & both the concentration and interactions between the elements, Bernstein , 1977). stainless steels for orthodontic purposes are of the 300 (austenitic) and 400 (martensitic and ferritic) series, (Thurow, 19721' The a most commonly used alloys in orthodontic practice are AlSl Type 302 and modified for welding AlSl Type 304, (Kusy et al, 1988)' These typically it have between 17-2Q per cent chromium to passivate the steel, (ie make ,stainless'), I - '12 per cent nickel to stabilise the austenite phase at room prevent by temperature, and a maximum of 0.15 per cent carbon to corros¡on the formation of unwanted carbides, (¡b¡d)'
An austenitic alloy generally has a 'face-centred-cubic' ('fcc') crystal lattice, 'body- although with hardening by cold working, some transformation to a centred-cubic' ('bcc') lattice can occur, (Greener et al, 19721' These (after lattice configurations are shown in FTGURES 2 - 1 and 2 - 2, Combe' 1975, p25). 16
--f I
I I
I t_ l_ I ---+ t,
I
,{ I { l_ I I 1---- I I
I _t I t--
F]GURE 2 - 1
'Face centred cubic' ('fcc') lattice'
FIGURE 2. 2
'Body centred cub¡c' ('bcc') lattice' 17
The detailed geometry of either an 'fcc' or a 'bcc' crystal lattice will determine the sites and possible concentrations of solutes such as interstitial carbon and nitrogen, as well as the diffusivity of alloying elements and, importantly, the metals' behaviour during plastic forming and manipulation' Carbon and nitrogen atoms enter an iron crystal lattice at a rate that is determined by the available interstitial space, but generally occupy the smaller octahedral rather in the than the larger tetrahedrat interstices, as this requires less strain energy movement of nearest neighbour iron atoms. Larger alloying elements such as chromium, manganese or nickel have an atomic size approaching that of the iron, and as such become part of the lattice by substitutional incorporation. The formation of plain carbon steel by the addition of carbon to iron has a great strengthening effect. since the solubility and diffusivity of carbon and nitrogen can be altered by heating, heat treatment can partly account for the precipitates strengthening of steels. on heating, excess carbon or nitrogen out of a saturated or supersaturated solid .solution, and thus in favourable may circumstances, can increase strength. Alternatively, extended heating result in an elimination of solute concentration gradients, ie an annealing homogenisation, (Honeycombe, 1981).
Martensitic alloys are formed by athermal transformation of austenite to a ,bcc' distorted lattice. They can be hardened by heat treatment in the same manner as a plain carbon steel. This procedure increases the hardness,
strength, and proportional limit, but decreases the corrosion resistance of the alloy. Martensitic alloys are restricted to a maximum of 18 per cent chromium, (Phillips, 1973). They have a higher edge-strength but lack the plasticity of austenite and generally fracture more easily. 18
Orthodontic wires are generally made from 18 - I (18 per cent chromium and 8 per cent nickel) austenitic stainless steel. ln these wires, strength and hardness increase with a reduction in diameter due to cold drawing when forming the wire. Drawing involves forcing the wire through a die by a tensile force applied at the exit side of the die, where most of the plastic flow is caused by compressive forces that arise from the interaction of the metal with the die. The process is generally carried out at room temperature, however, due to the large deformations that occur in the wire, a marked rise in temperature occurs during drawing, (Dieter, 1961).
Wire drawing begins with hot-rolled wire rod which is cleaned by pickling to remove any scale that may cause later defects, either due to inclusions or surface contamination. The wire is then plated with a thin layer of copper or tin for wet drawing, or coated with lime for dry drawing. ln wet drawing, the entire die is immersed in a lubricant of fermenting rye-meal liquor or alkaline soap solution. ln dry drawing, the added lime acts aS an absorber and carrier of the lubricant, and further serves to neutralise any remaining acid from the pickling process. tn dry drawing, the lubricant is usually grease or soap powder, (¡b¡d).
For fine wire, such as ofthodontic wire, a number of draws are necessary through a series of dies with progressively decreasing diameters. To begin drawing, the end of the wire is pointed, often by electropolishing, passed through the die, and attached to ä draw block, or capstan, that ¡s part of a drawing drum, and then 'pulled' by either a chain drive or hydraulic mechanism. Drawing speeds may vary from ten to one hundred metres per minute, depending on a number of variables, including the character¡st¡cs of the die, the type of wire being drawn, or the physical property requirements of 19 the finished product. Drawing force, which is the sum of the force required to decrease the diameter uniformly, (ie tensile elongation), the force to produce non-uniform shear deformation of the surface layers at the entry to and exit from the die, (ie redundant work), and the force required to overcome friction between the wire and the die wall. The required force depends on the die angle, the percentage reduction, the flow stress of the material, and the die friction, which is dependent on the die material, the lubrication, and the drawing speed, (¡b¡d).
Most dies are made from tungsten carbide to allow long die life. A typical die (shown at FIGURE 2 - 3), has a truncated conical shape with an entrance angle, which is large enough for the undrawn wire and the lubr¡cant that adheres to the die. The approach angle (or'half die angle': 'ctr') is that part of the die where the actual reduction in wire diameter occurs. The bearing surface of the die acts as a guide for the wire as it exits the die.
Variation in the value of 'cr' will alter the homogeneity of the surface of the drawn wire. For a given wire diameter, the amount of shear deformation in the direction opposite to the draw pull increases with increasing half die angle. Only the central axis of the wire undergoes pure elongation. "For large half die angles, the large shear deformation results in an increased longitudinal tensile stress at the centre of the wire, that can exceed the fracture stress, and can thereby result in a cupping type of failure. For a given half die angle, the shear deformation becomes less important with increasing percentage reductions. In addition, because of increased surface shear deformation, the yield and tensile strengths of drawn wires are higher for larger half die angles. This effect is greater for lower reductions", (¡b¡d). 20
Die holder Drow heod
Die
Ø Jow
Beoring su rfo ce Entronce ,7 ongle
Approoch ongle \
FIGURE 2 - 3
Typical wire drawing die.
(Dieter, 1 961) 21
Non-uniform shear deformation is lower with smaller half die angles, but wall friction is greater, which suggests an optimal die angle for minimum drawing force for a given reduction. This is due to the ¡nteraction of half die angle, reduction percentage, flow stresses, the type of material, lubrication, and friction. All else being equal, the optimal half die angle increases with the amount of reduction. The drawing speed has little effect on the drawing force, but may profoundly affect the temperature, making lubrication more difficult. Another consideration is the 'back-pull' in the opposite direction to the drawing pull. This can arise directly from frictional forces in a poorly lubricated system, or might be deliberately applied in order to reduce die wall pressure and therefore frictional forces, enhancing die longevity' An increased back-pull also has the effect of decreasing the maximum possible reduction in diameter, (ibid).
As a result of cold drawing, orthodontic wire attains anisotropy because of the re-orientation of the individual grains. Two distinct types of residual stress patterns are seen, depending on the amount of reduction. "where the reduction per pass is less than one per cent, the longitudinal residual stresses are compressive at the surface and tensile at the axis, the radial stresses are tensile at the axis and drop off to zero at the free surface, while the circumferential stresses follow the same trend as the longitudinal residual stresses". This pattern is consistent with localisation of deformation in the surface layers of the drawn wire. For larger reductions, the residual stress d¡stribution is completely opposite to those reductions of less than one per cent. "The longitudinal stresses are tensile at the surface and compressive at the axis, the radial stresses are compressive at the axis, and the circumferential stresses follow the same pattern as the longitudinal stresses". It has been shown that for a given reduction, the longitudinal residual stress 22
increases as the half die angle increases. Maximum values of residual stress in cold drawn brass wires are obtained with reductions of about fifteen to thirty-five per cent, (ibid).
Much of the theory of wire drawing is considered as occurring with a material that exhibits pure plastic deformation, without friction. Even with the best tubricated die drawing system, this is unrealistic. These assumptions ignore the presence of friction and shear (redundant) deformation, or, only consider an ideal situation. No theoretical analyses based on slip field theory have grid been applied to wire drawing, and, since most theory is calculated with a system in two planes, such theories have only limited application to the actual process, (¡b¡d).
During manufacture, heating and cold drawing is done alternately until the \ heating at specified j wire attains its desired propert¡es. The amount of
I I intervals during diameter reduction depends on several factors, including the
diameter of the raw material, and the diameter of the finished product' The
use of heat is used to relieve stresses in cold worked austenite to improve the tensile elastic properties, such that the proof stress, or yield point is improved' This improvement ranges from ten to twenty per cent compared to untreated wires and allows ease of forming for orthodontic purposes, (Begg' 1954)'
Heat treating a strain hardened austenitic alloy can lead to annealing (recovery or recrystallisation), (Phillips, 1973), if temperatures above 500'C are used. Unfortunately this occurs within the soldering temperature range of sitver solder, (from 650 to 750"c), such that careless soldering of an orthodontic wire can tead to a loss of strength, but in particular a loss of stiffness and springback, (Wilkinson, 1962). To this end, silver solders r 23 with lower melting ranges are used where possible, since this has a greater potential for reducing wire annealing than the actual time of heat exposure, (wilkinson, 1963). welding, even though a higher temperature is involved, localises the heat effect to the area being welded between the joints electrodes and reduces gross c'hanges in properties. However welded do have decreased corrosion resistance, (Phillips' 1973).
lf Stainless steel wires are subjected to lower temperatures over time, in the range 260 to 460"C for 0.5 to 75 minutes, the wire can be tempered such that an increase of up to 25 per cent in the (tensile) strength and in the relative stiffness occurs, (wilkinson, 1962). Heating above 400'c will reduce the stainless steel wires' resistance to corrosion since, due to the formation of chromium carbide at the grain boundaries, the proportion of chromium diminishes in the solid solution adjacent to these boundaries, (Gardiner & Aamodt, 1969). The process of heat treating stainless steel wires is a frequent feature in the manufacture of orthodontic wires since it allows large reductions in diameter to produce wires with qualities of improved strength
and stiffness.
An improvement in the elastic properties of stainless steel wires occurs during heat treatment after manufacture, (Funk, 1951; Howe et al, 1968; Brenner et at, 1981). Bass & stephens, (1971), showed that for low temperature heat treatment, there was no improvement in the flexibility of finger springs, although they exhibited an increased resistance to fatigue' Detgado & Anderson, (1963), suggested that the effect of low for temperature heat treatment on tensile strength and on a wire's capacity bending is clinically insignificant. 24
Any improvement in the elastic properties of stiffness and springback resulting from heat treatment, may be accompanied by a concomitant loss of formability or ductility, (Waters et al, 1976; Craig, 1978). Severe cold work during drawing leads to the production of martensite from 'deformation induced' metastable austenite. The martensite may be subsequently transformed back to austenite during 'stress relief' heat treatment, (Khier et al, 1988).
The relative advantages and disadvantages of the heat treatment of 18:8 stainless steel wires have been summarised by Waters, (1976)' as follows:
Advantages: o "Less internal stress in the wire, hence
(i) more uniform ProPerties
(ii) better fatigue ProPerties
a Slightly improved elastic properties.
Disadvantages: a Takes time and trouble,
a Suitable heating apparatus required,
a Elastic recoil of approximately 0.2 per cent occurs which necessitates further adjustment of the appliance,
a More susceptible to fracture if subject to repeated
adjustment,
a Corrosion resistance reduced,
a Poor colour."
On balance, it would appear that the heat treatment of high tensile wires is of insufficient clinical value to warrant its routine use. 25
The properties of orthodontic stainless steel wires are well documented in the work of squires, (1991), where much of the fundamental metallurgy of Honeycombe, (1981), cited from his text, 'Steels, microstructure and properties, first edition', is coupled with the earlier work of Colombler and Hochmann, (1967), in their text, 'stainless and heat resisting steels, first edition'.
statistics show that gg per cent of orthodontic practitioners continue to use stainless steel wires. The wires are both dependable and cheap, suggestive of continued use well into the next century, (Kusy et al, 1988). 26
The purpose of standards in the manufacture of any product is to promote and ensure quality. ln developing and adopting a standard, applicability of the standard to actual product use is important. The first standard for orthodont¡c wires was introduced by Paffenbarger and associates in 1932, (Nikolal et al, igBS). The Australian Standard, 4S1964, for 'Resilient orthodontic wires (excluding precious metal wires)' was formerly published in 1965 as ,Australian Standard T32'. This temporary standard had been originally developed from the then existing British Standard. Following discussions between interested scientific, industrial, and governmental organisations, a new standard was published in 1977, (Standards Association of Australia, 1977\.
The latest version of the American Specification, No. 32,lor'Ofthodontic wires not containing precious metals' was accepted and published in 1977, (Nikolai et â1, 1988).
The British standard, Bs 3507 : 1976, 'orthodontic wire and tape and dental ligature wire' was published in 1976, (British Standards lnstitution' r 976).
The basis of both the Australian and American standards is to classify wires according to their tensile properties. The Australian Standard specifies 6 types of austenitic chromium - nickel stainless steel wire according to TABLE 2 - 1. ln comparison, the American specification is for 2 types of wire not conta¡ning precious metal, according to TABLE 2'2. By contrast, the British Specification suggests the suitability of Grade 302S25 (AlSl Type 302) wire' and considers tensile data in relation to wire diameter. Unlike other 27
Tensile Strength, MPa TYPE Minimum Maximum
1 1720 2240
2 2241 2400
3 2401 2550 4 2551 2700 5 2701 2860 6 2870
TABLE 2 - 1
Classification of wires (by strength) as required by the Standards Association of Australia, AS 1964 - 1977 lor Orthodontic wires (resilient) (excluding precious metal wires)
Flexure Yield Strength - MPa TYPE (2.9o offset) Minimum Maximum
1 700 2400 il 2500
TABLE 2 - 2
Yield strength requirements as required by the American Dental Association Specification Number 32lor orthodontic wires not containing precious alloys. 28 standards, the British Specification provides additional requirements for multi- stranded wires and for soft ligature wires.
Tests are performed on the wires to determine their mechanical properties. The Australian Standard requires three tests; tensile strength, wrapping ability, and resistance to bending. The American Specification requires tests for the propefties of flexure yield strength, modulus of elasticity, (known in the published standard as the 'modulus of stiffness and sometimes referred to as the 'modulus of rigidity'), and resistance to bending. The British Specification requires tests on extra hard drawn wire to determine the tensile strength and resistance to failure on bending as well as requiring the wire to be finished to a clean smooth finish. ln essence then, these three standards are testing similar wire properties, particularly by tensile and bending tests.
The 1ensile test' (AS 1964) records the force applied to fracture the wire being held by two grips. However, the 'flexure yield strength test' (No. 32) derives the yield stress from cantilever beam tensile tests, and goes further by requiring a mathematical derivation of Young's modulus of elasticity ('E'). Goldberg, (1983b), suggests that the use of tensile data to determine flexural data for fine orthodontic wires is clinically irrelevant. The springback of orthodontic wires is affected by many factors, including the wire size and shape, the modulus of elasticity of the material, and the design of the appliance. There appears to be no consensus as to the best orthodontic definition of 'elastic limit'. The usual method for determining elastic limit is to stress a cylindrical specimen in tension and then determine the point of flexure by plotting a stress-strain diagram. This method is considered inappropriate because it tests the material in tension, whereas orthodontists primarily stress wire in bending. Such tests have been suggested as being 29 obscure to the clinician because they d¡ctate quantifications of mechanical, rather than structural properties, (Nikotai et al, 1988).
Deter¡nining the mechanical properties of orthodontic wires under bending conditions is preferable to using tensile tests, since bending is representative of clinica! conditions, (Asgharnia & Branttey, 19SO) The American Specification requires a cantilever test that uses a one inch, (25.4 mm) length of wire. This length has been used by Drake et â1, (1982)' The cantilever test has been shown to be valid for wires with a diameter greater than 0.7S millimetres. Serious short-comings were noted with the cantilever bend test when determining values for mechanical properties for fine wires, attributable to the test specifications, (Asgharnia & Brantley, 1986)'
Branttey et at, (1978), used beam theory techniques, to show results for the measurement of varying lengths of the same material which gave 'E' values that differed by more than 100 per cent, between half, one, and two inches, (12.7,25.4,50.8 mm) specimens. lt was suggested that to determine the physical properties of wires, such as Young's modulus of elasticity ('E') using very short lengths of wire in a cantilever bending test, (Brantley & Myers, 1979).
The British Specification seeks tensile testing in accordance with British specification Bs4545. Details of this specification are not contained in the ,Dental standard', making comparisons difficult. lt ¡s likely, however, that the testing is similar to the Australian and American tests.
The use of mandrels with varying diameters, to wrap wire and assess the degree of deformation under a given load, has been suggested as being 30 clinically relevant in assessing the springback of orthodontic wires, (stephens & waters, 1971; Waters, 1981; lngram et â1, 1986). The authors suggest that the wrapping test should supersede the tensile test since it has greater clinical applicability, particularly where a wire needs to be deflected to engage an orthodontic attachment. This is supported by williams, (1974). The wrapping ability test (AS 1964) is performed using a consistently rotating mandrel to form closed coils of wire that are microscopically examined for defects. The coiled wire is then 'pulled out'to five times its length and the wire re-examined for defects.
The,resistance to bending tests'are similar; AS 1964 requiring successive 80 degree bends, No. 32 requiring successive 90 degree bends, and BS 3507 requires 40 degree bends. The number of cycles to fracture is reported.
A three point and a four point bending test have been used by Kusy & Dilley, (1gg4). These tests were for the assessment of Young's modulus in mult¡-stranded wires. tn supporting a wire between two points and deflecting it with another one or two points, these tests are considered more clinically point relevant compared with a cantilever bend test. These three and four bending tests are considered to simulate the stress state in an orthodontic wire, as well as eliminating the gripping of a wire in a two point cantilever test' The three point bending test is considered simpler in terms of geometry, and thus makes for easier results calculation' (¡b¡d).
An alternative five point bending test to determine the ability to engage a wire to a displaced tooth has been suggested, (Nikolai et al, 1988), because of problems identified in relating a single ended, 'in-plane' couple loading' to an actual clinical activation. Load-deformation information is converted to beam 31 theory data, and physical propefties derived. The use of a five point bend test is considered superior as it mimics the clinical process of engaging a wire into adjacent brackets. This test is considered particularly appropriate for the newer nickel-titanium, B-titanium and near-cr-titanium wires, although no allowances for differences between oral and 'bench' temperatures have been included, (Schultz, 1988).
The orthodontist is faced with such a large selection of wires, that choice is based on either'feel', clinical impression, (Hazel et al, 1984)' and/or standard physical test data, which, are mostly irrelevant, despite their scientific nature, (Thurow, 1972). Factors such as wire curvature, bracket angulation, bracket friction, method of wire ligation, inter-bracket distance, and the effects of occlusion will alter the force systems in a wire, (Adams et al, 1987; Burstone & Koenig, 1974; Moran, 1987; Rock & wilson' 1988; Waters et al, 1975a, 1975b). 32
ln the Begg orthodontic technique, perhaps the most important bends one can place in an arch-wire are the anchorage bends. From these bends, the appliance derives its ability to open the bite and address anchorage concerns. lndeed, the ability to perform successful Begg treatment is based on opt¡mal anchorage control, (Sims, 1971). Simitarly, with any orthodontic appliance, effective clinical control of the appliance is achieved by ,,understanding the forces generated when the arch-wire is engaged to the attachments on the teeth", (Sims, 1977). The application of these forces and their assoc¡ated moments and couples is often poorly understood, (Mulligan, 1979). The simplest of arch-wire manipulations, a single bend, can introduce a complex interaction of forces once the arch-wire is placed in the mouth. lndeed, at any instant in time, an elastically deformed arch-wire has an internal stress system equilibrated externally by the constraints imposed by orlhodontic attachments. The equilibration therefore, defines the force system acting on the teeth, (Borghes¡' 1973).
Bending is the change that occurs when a straight length of material (metal) is changed to a curved length. ln beam theory terms, this occurs when a transverse load is applied to an object, the result of which is a bending moment comprising the sum of all moments applied. The stresses accompanying the bending moment are the bending stresses comprising compressive and tensile stresses in equilibrium. Bending is generally accompanied by transverse shear and torsional shear stresses, (Twelftree' 197 41. plastic deformation, or bending, occurs by linear disturbances ¡n the atomic arrangement, or dislocations, that move through the crystal lattice. 33
Dislocations move through the lattice until they emerge at a free surface, or until they meet a lattice discontinuity, such as, a point defect, an immobile dislocation caused by the interaction of other dislocations, a foreign atom or precipitate particle, or, a grain boundary. ln polycrystalline materials, dislocations tend to accumulate at grain boundaries and contribute to work hardening, (Phillips, 1973). Work hardening is the requirement for the application of increased stress to deform grains. This occurs following an increase in the number of dislocations within the grains that become tangled, and, the increase in dislocation density at grain boundaries. Actual grain size can have an effect on the yield stress, since for larger grain sizes, and therefore larger grain surface areas, the accumulation of dislocations is greater. These in turn cause higher stress concentrations in neighbouring grains. tn practical terms, the finer the grain size, the higher the resulting yield stress, (Honeycombe, 1981). To this end, the production of high tensile wires partly involves the manipulation of mechanical propert¡es by alterations in the microstructure, or grain size, by a balance of cold work during drawing, and intermediate heat treatment between successive draws.
Where plastic deformation occurs at temperatures below the recrystallisation temperature, the process is known as 'cold working'. Due to the strain hardening effect, there is a commensurate increase in the hardness, yield point, and strength after cold work, as well as a decrease in the metal's ductility, (¡b¡d).
When a wire is bent, the resulting bend radius, 'R' is the radius of the concave surface; the neutral axis is the circumferent¡al f¡bre at which the strain passes through zero. ln plastic bending the neutral axis does not remain at the half thickness, as occurs in elastic bending. For a sharp bend, the neutral axis is 34 closer to the inside of the bend, usually assumed to be at 0.45 times the thickness. When sheet metal is bent, the final length increases because the thickness decreases. As the bend radius reduces, so the thickness decreases. The changed length of the centre line of a bent specimen is known as the bend allowance, 'B', and is useful in determining the required length of a blank prior to a bending process. The neutral axis is assumed to have a radius of curvature of:
R + 0'45.t
where: R : bend radius,
t:t- thickness.
The bend allowance 'B'can therefore be calculated by
B = (R+0.45.t )#
where: B= bend allowance, R= bend radius, t: thickness, (Dieter' 1961). cx, = bend angle in degrees,
The above terms and definitions are illustrated in FIGURE 2 - 4.
Given the small diameter of orthodontic wires, it is unlikely that the tolerances of bend allowance are clinically significant. The aforementioned theory however, in some part helps to give a greater understanding of the dynamics of orthodontic wire bending. 35
t A 0.45 t +
R
FIGURE 2 - 4
lllustration of terms used in the theory of plastic bending.
where: B = bend allowance, R = bend rad¡us, | = thickness, cr = bend angle in degrees, (Dieter' 1961). 36
More important is a consideration of the strains inherent in bending. These are circumferential, transverse, and radial stra¡n. Of greatest significance is strain in the tongitudinal or circumferential direction. This strain is critically dependent on the type of bending process. Three point bending for example, produces quite a non-uniform distribution of strain, whereas bending by wiping or wrapping allows for a progressive increase in bend angle and thus some gradual change in strain. Near complete uniformity in circumferential strain however, does not occur in bends less than about ninety degrees. ln bending round wires, the transverse strain is important, since the 'edge' of the bent specimen is subjected to uniaxial tension, given that stress normal to a free surface is always zero. Thus, because of 'Poisson's effect', a transverse compressive strain results. 'Poisson's effect' relates the ratio of strain in the transverse direction relative to the strain in the longitudinal direction. Poisson's ratio for a perfectly isotropic elastic material is 0.25, but for most unformed metals is closer to 0.33, (¡b¡d).
The theory of bending suggests that the strain Increases with decreasing radius of curuature, thus if a theoretical model ¡s established where change in thickness is ignored, the neutral axis will remain at the centre of the specimen, and the circumferential stretch on the 'top'convex surface will be equivalent to the circumferential compression on the 'bottom'concave surface. Where the change in thickness is considered in experimentation, the circumferent¡al strain on the tension surface is considerably higher than the compression surface. The strain on the compression surface approaches that of the theoretical model. For any given material, the bend radius has a minimum possible value, since for a smaller value, the outer tensile surface will fail, generally by cracking. This minimum bend radius varies between materials, always increases when the material is cold worked, and is smaller for more 37 ductile materials. The minimum bend radius however, is not a precise value, since it depends on the geometry of the forming, or bending process. The minimum bend radius can be increased slightly by polishing the specimen, since this reduces the ability of grain dislocations to concentrate at 'edge deformities', (ibid).
Springback can be a problem in bending, since it represents the dimensional change occurring after the forming tool has been removed. lt results from the changes in strain from elastic recovery, resulting in a reduced total strain, and is related to the bend radius and the wire diameter. The springback ratio, 'K', is defined as:
ctt K cre
where K= springback ratio 0o= bend angle before release of load,
Clf = bend angle after release of load, Ro= bend radius before release of load, Rf= bend radius after release of load, d= wire diameter
For a given material, the elastic recovery, or springback, is greater for higher yield stress, (which can be altered by heat treatment), and is greater for a lower elastic modulus and increased amount of plastic strain. The easiest way to compensate for forming springback is to over-bend the specimen, (¡b¡d). 38
Stainless steel orthodontic wires have a range of ultimate tensile strengths. The dynamic response of metals is altered by 'size effect' or diameter and length. Detrimental effects from cold drawing inctude surface geometric irregularities that often lead to failure after use or cyclic fatigue, (Coquillet et ât, 1g7g). The elastic limit of almost all structural material, including stainless steel, is lower in tension than it is in compression. Therefore, when a bend is made in,a wire, permanent deformation wlll occur firstly on the tension side, (lngram et al, 1986). For wires with an extremely high yield strength, a single bend, such as an anchorage bend in the Begg technique, (Begg, 1956), could lead to failure or fracture if the wire is poorly handled, (Witcock, 1989).
The study of stress relaxation in high tensile wires with simulated Begg anchorage bends has been carried out by Hazel et al, (1984). They concluded that 0.016 inch (0.4064 mm) Wilcockl wires undergo significantly less stress relaxation than other stainless steel wires. This supports the work of Twelftree, (1974), who compared the stress relaxation properties of Wilcock wires against Unisil2, Dentaurum3, and Elgiloy+ wires.
A strain gauge analysis of continuous arch-wire mechanics produced by a contraction utility arch, as used by proponents of the bioprogressive technique, (Ricketts et al, 1979), has been pioneered by White et al, (1979), and subsequently refined and reported, (Murphy et al' 1982). A disparity in the amount of force delivered to different anterior teeth was found to be a factor of arch form. The lateral incisors were found to sustain the greatest amount of applied force under all conditions. Greater anterior
1 R.¡. Wilcock Scientific and Engineering, Whittlesea, V¡ctorla, Australia. 2 gtt¡ / Unitek Corporation, Monrovia, California, USA' 3 Dentaurum, Pforzheim, Germany 4 Rocky Mountain Orthodontics, Denver, Colorado, USA. 39 contract¡on forces were found in a narrow tapered arch form when compared to an ovoid arch form. The two different arch forms were considered to represent clinical extremes, (¡b¡d).
Of particular interest, a number of authors have used a jig to accommodate an average arch form, in order to measure the anterior intrusive force exerted by the tip-back bend or anchorage bend mechanism. White et al, (1979) and Murphy et à1, (1982), were able to measure arch-wire force components with a series of strain gauges. However, most subsequent workers have measured forces at the central point of the arch, due mainly to ease of technique and cost. Excessive friction was noted by Hazel et al' (1g84), when measuring arch-wire forces in the canine region with a bar mechanism. Consequently, no conclusion could be drawn as to forces existing in other areas of the arch, such as in the lateral incisor, canine, or premolar regions.
Similar models have been used to measure the intrusive forces of arch-wires from a single point at the anterior centre of the arch-wire, (Borghesi' 1973; sims, 1977; Thornton & Nikotai, 1981; Rohan, 1982; Willmot' 1983 and Sampson, 1992). The initial forces are summarised and compared in TABLES 2 - g and 2 - 4, and represented in graphical form in FIGURE 2 - S. Prior to these archform investigations, only simple two dimensional beam theory experiments had been performed and documented, (Borghesi, 1973). 40
for different wire grades - grams Author and Measured intrusive force tip-back bend Regular Special Specialplus Premium plus Borghesi :45o 36.0 Thornton&Nikolai :45o 46 Willmot:45o 73 68.0 74.0 84.1 93.2 Sims :45o 70.0 80.0 Sampson :40o 32 42 47
TABLE 2 - 3
Anterior intrusive force from anchorage bends placed in 0.016 inch (0.4064 mm) Wilcock wires as measured by four separate authors.
Of note is that the values obtained by Hazel et al, (198a) in TABLE 2- 3 have been cited incorrectly by Rock & Wilson' (1988). They suggest that an intrusive force of 195 grams, rather than up to 93 grams, was measured. 41
bends - grams Author Measured intrusive force for different anchorage 100 15" 20" 300 40" 45" 50" 60"
Borghesi,'73 (Note 1) 7 12 24 32 36 41 Sims,'77 15 30 50 Thornton&Nikolai,'81 28 46 62 Rohan,'82 36 54 71 85 94 Willmot, '83 (Note 2) 10 18 26 44 67 74 79 100 Hazelet al,'84 84 Sampson,'92 42
TABLE 2. 4
Anterior intrusive force from anchorage bends of varying degree placed in 0.016 inch (0.4064 mm) 'special plus' wilcock wire as measured bY several authors.
Note 1: Wire used by this author was 'special'grade. Note 2: lnformation derived from graphical representation in cited literature. 42
100
90 .) BO
70 I Borghesi f -#Sims 60 / & Nikolai Force *Thornton 50 t (grams) 4Rohan j (. 40 Willmot r #Hazel et al 30 { I Sampson 20 ) I 10 I 0 0 10 20 30 40 s0 60 Anchorage bend (degrees)
FIGURE 2.5
Graphical representation of variation in intrusive force as a function of degree of anchorage bend as measured bY several authors.
The most important feature of the graph¡cal representation is the linear, or near linear, behaviour of w¡res in relation to the degree of anchorage bend. Small differences may be attributable to such things as rolling of the arch-w¡re
in the molar buccal tube in the transverse plane, the effects of friction between the wire and the molar buccal tube or distal stops, or, inconsistencies in the distance of the anchorage bend from the molar buccal tube. 43
Several authors have measured and recorded the difference in intrusive force ,l progressively moved away from the r'ì that occurs as the anchorage bend is molar buccal tube. These dynamic mechanics are a feature of Stage 1 of Begg mechanotherapy, (Sims, 1975). As can be seen from TABLE 2 - 5, as the anchorage bend moves closer to the molar buccal tube, the anterlor intrusive force increases. Thus, if the requirement for anterior intrusive force decreases, or remains the same as anterior tooth retraction occurs, (and the anchorage bend moves closer to the molar buccal tube), then the degree of anchorage bend must be reduced.
Distance of anchorage bend anterior to buccal tube 0mm 1mm 2mm 3mm 4mm 5mm 6mm 7mm 8mm
Borghesi, '73 36 't |f Sims,'77 60 40 35 25
i Thornton & Nikolai, '81 70 63
Rohan, '82 (Note 1) B9 85 82 79 76 73 Willmot,'83 53 50 47 43 40 36 33 30 Sampson,'92 (Note 2) 36 35 32
TABLE 2.5
Anterior intrusive force generated by 0.016 inch (0.4064 mm) Wilcock'special plus' arch-wires with 45 degree anchorage bends as a function of distance of anchorage bend from the buccal tube as measured bY several authors.
Note 1: Anchorage bend of 50 degrees used by this author. Note 2: Anchorage bend of 40 degrees in 'special' grade wire used by this author.
I 44
Willmot, (1983), showed that a near linear relationship exists between the distance from the molar tube of a 30 degree bend, and the measured anterior intrusive force exefied by a 0.016 inch (0.4064 mm) arch-wire. For each additional millimetre the bend was moved in front of the molar tube, the intrusive force reduced by approximately 6.6 per cent. However, other differences in experimental design prevent greater correlation of results between different studies. The similarly linear findings of the various authors from TABLE 2 - 5 are represented in graphical form in FIGURE 2 - 6.
90 \ )- BO \ 70 .- 60 \ Borghesi --l-Sims -l 50 lt Force Thornton & Nikolai i (grams) ; OO êRohan Willmot 30 ,..:- I Sampson 20
10
0 01 234s6 7B Distance form molar tube (mm)
FIGURE 2 - 6
Graphical representation of var¡ation in intrusive force I as a function of distance of a 45 degree anchorage bend from the buccal I molar tube as measured by several authors.
r 45
The 'ideal' position of the anchorage bend has been suggested as being "approximalely 2 mm mesially to the buccat tubes", (Batdridge, 1973), and is without doubt contingent upon inter-related therapeutic variables as dictated by clinical diagnosis, (Sims, 1975).
The use of bends and force production in arch-wires has been studied and reported oî, (Mulligan, 1979; Burstone & Koenig, 1988; Ronay' 1989; Bequain, 1992). Most of the work of these authors has been done in two dimensions or with a straight piece of wire. However, one of the consistent points elicited by their work, has been that in overcoming the required wire deflections during the initial placement of an arch-wire, a variety of flexures involving a mixture of torsional, shear, and flexural stresses are involved.
Those mechanical properties considered important for orthodontic wires have been used in an attempt to understand the forces involved in such a situation, (Waters et ât, 1975a, 1975b, 1981, 1987; Burstone & Koenig' 1976; Waters, 1976; Lane & Nikolai, 1980; Goldberg et al, 1981; Drake et at, 1982; Hazel & West, 1986; Schaus & Nikolai, 1986; Gullotta et â1, 1987; Cavina & Waters, 1988; Johnson & Lee, 1989; Koenig & Burstone, 1989; Demange, 1990; Van den Bulcke, 1990). The main findings of these papers are beyond the scope of this project.
An attempt has been made to simplify the inter-relationship of many arch-wire properties and arch-wire types by the use of simple line diagrams, or 'nomograms' that are derived from the mathematical relationship of variables to one another, (KusY, 1983). 46
It is important to remember that in the application of a theoretical elastic model to orthodontic arch-wires, accurate prediction of bending moments for strictly elastic behaviour can be made. However, when even the smallest permanent angular deformation occurs, application of this theoretical elastic model is spurious, (Goldberg, Morton & Burstone' 1983).
On the basis of the information presented in this chapter, it was decided to conduct an investigation into the effects of different pl¡ers on bending high tensile wires. Additionally it was decided to look systematically at the 'tensile test' as defined in the Australian Standard, and in particular at the effects of ageing on tensile data. Finally, the force levels associated with a typical Begg stage 1 arch-wire were poorly understood in the literature, and an evaluation was undertaken of anchorage bend effect on anterior intrusive force levels. 47
CHAPTER 3
MATERIALS AND METHODS
PREPARATION OF WIRES FOR SEM
MATERIALS lnvestigations into the behaviour of orthodontic wires incident to bending have been undertaken with two spools of A.J. Wilcock'premium Plus's wire, batch ,TT-TAT', (each of 7.6 metres) with a nominal diameter of 0.016 inches and an actual diameter of 0.0159 inches (0.4064 / 0.4039 millimetres).
The 'raw' AlSl Type 302 wire was purchased by A.J. Wilcock as an 8'85 kilogram spool of 18:8 stainless steel wire, from a supplier in Europe, and then treated and drawn to the parameters detailed in TABLE 3 - 1' During processing, the wire was coated with a lubricant prior to drawing' This lubricant was then removed by washing in industrial solvents prior to any
subsequent heat treatment and further lubrication and drawing, or, packaging. The batch of wire was 'finished' on 07 December 1983. Further manufacturing details cannot be divulged as they are bound by the dictates of "Commercial - in Confidence".
Processing of the wire is illustrated in FIGURE 3 - 1.
5 A.J. Wilcock Scientific and Engineering Equipment, Whittlesea, Victoria, Australia. 48
t \[- | , ¡\{,--. ¡ I t. I= a- - I ¡ F'l t a t .4
-'.-S¡\rì I . n l-g-; I
rv¡Ètlff i Irh-; I I :lt' O
I
FIGURE 3.1.
Wire drawing by A.J. Wilcock Scientific and Engineering Equipment (with kind permission from Mr. A.J. Wilcock) 49
Ultimate tensile Diameter strength
lnches mm (Pounds) 103
0.0159 0.4039 90 3.1
TABLE 3.1
Diameter and Tensile properties of batch'TT-TAT'of Wilcock 0.016 inch, (0.4046 mm), 'premium plus'
Most finished batch spools are of the order of 25,000 feet in length (7,620 metres). Prior to cutting and packag¡ng, both ends of the batch spool of wire are tensile tested to determine its classification. This is done with a 'bucket test'. One end of the wire is wrapped around a one inch, (25.4 mm), diameter round pipe, and fastened with a grub screw. The whole assembly is supported in a large engineer's vice. The other end is wrapped around the handle complex of a metal bucket of known weight. lmperial pound weights are progressively added slowly and with great care, to avoid snapping the wire, until the wire fails. lf the wire fails between the two anchor points, the test is regarded as valid, and one pound is subtracted from the tota¡ weight in imperial pounds, to give the breaking strain of the wire. lf the wire fails at one of the anchor points, the test is repeated. Given that the wire's diameter is known, the tensile strength of the wire can be determined. The manufacturer
claims that a'dead weight'test is the most accurate method of measuring the
ultimate tensile strength of high tensile wires, (Wilcock' 1994). 50
Calculation is achieved with the following equation
+2240 rr=+9n.d¿
where: TT= strength in imperial 'tons tensile', BS= breaking strain in imperial pounds, d= wire diameter in imperial inches.
In addition, the wire was subjected to a wrapping test every two thousand feet (609.6 metres). This is performed by wrapping the wire at least five t¡mes around a mandrel of the wire's own diameter and visually checking the sample for any signs of imperfection such as cracking or splitting.
Details of the normal composition of AtSl Type 302 stainless steels are shown in TABLE g - Z. These details are compared with the composition of 'Australian' wire6, (probably 'special' or 'special plus' grade, but not cited), determined by Newman, (1963). These alloys should contain 17 lo 19 per cent chromium to passivate the steel, that is make it 'stainless', I to10 per cent nickel to stabilise the austenitic phase at room temperature, and a maximum of 0.15 per cent carbon to prevent the formation of carbides, (Kusy et al'
1 e88).
Given the presence of molybdenum, the wire tested by Newman' (1963)' cannot be classified as an AlSl Type 302 stainless steel wire. Howe et al, (1968) suggest that AlSl Type 316 grade wire has superior elastic properties compared with AlSl Type 302 wire. lt contains up to two per cent
6 Manufactured by A.J. Wilcock, Australia, marketed by TP Orthodontics, lnc., La Porte, lndiana, USA. 51 molybdenum to increase the wire's corrosion resistance, however, there is also an additional two per cent nickel. Thus, there is difficulty in classifying the wire tested by Newman, (1963), according to published tables of AlSl Types, (ASTM, f 99O). lnformation from the manufacturer suggests that the wire stock purchased for processing is now an AlSl Type 302, (W¡lcock,
1 994).
Type 302 comPosition, Per cent Element Atsl- 19æ Wilcockwire (ASTM, 1990) (Newman, 1963)
Carbon 0.15 (maximum) 0.095 Manganese 2.00 (maximum) '1.17 Phosphorus 0.045 (maximum) 0.022 Sulphur 0.030 (maximum) 0.032 Silicon 1.00 (maximum) 0.48 Molybdenum 0.21 Nitrogen 0.10 (maximum)
Chromium 17.00 - 19.00 18.21 Nickel 8.00 - 10.00 8.97
TABLE 3 - 2
Normal elemental comPosition for Alsl Type 302 stainless steel, (ASTM, 1990)' compared to 'Australian' wire as determined bY Newman, (1963). 52
M ETHO DS
Short lengths of wire were removed from the spool and initially straightened by running the wire through the fingers prior to being cut into 30 millimetre lengths. A bend of known degree was then placed in each segment of wire. This was done using a bend guide pattern similar to that in FIGURE 3 - 2.
Bends were made using pliers as detailed in TABLE 3 - 3. The different plier beaks are shown in detail in FIGURES 3 - 3 to 3 - 6. The high tensile pliers, (after Mollenhauer), were used to test selected specimens with larger degree bends following analysis of results obtained with the other pliers. This is discussed further in ChaPter 4.
prePared Pliers used to Specimens prepare specimens 100 200 30" 40" 50' 600
Light wire - s/s beak 7 Tweed torquing J
Light wire - ./ Tungsten carbide beak
High tensile wire - ,/ J Tungsten carbide beak 19
TABLE 3 - 3
Details of specimens PrePared with various plier tYPes and magnitude of bend Placed.
7 Catalogue number 5162, Crovana, Germany. I Catalolue number 442, Rocky Mountain Orthodontics, Denver, Colorado, USA' I Catalogue number 105, Dentronix, lvyland Pennsylvania, U.S.A. 10 Catalogue number AW100-240, A.J. Wilcock Scientific and Engineering Equipment' Wh¡ttlesea, Victoria, Australia. 53
90 EO 70 c0 !i0
40
30
20
10
0
ANCHORAGE BEND
FIGURE 3 - 2
Anchorage bend forming guide. 54
I
I i
I
;
L
i I I
ì I I
:
rì
ìi
!i t,j ^l ;l
I
,ì (.,L
¡
FIGURE 3. 3
Detail of beaks of stainless steel tipped light wire pliers.
t
f 55
FIGURE 3 - 4
Detail of beaks of steel tipped Tweed torquing pliers I stainless
þ 56
FIGURE 3 - 5
Detail of beaks of tungsten carbide tipped light wire pliers { 57
FIGURE 3 . 6
Detail of beaks of tungsten carbide tipped high tensile wire pliers. 58
Each bend was placed with a single bending action to minimise the introduction of a strain history in the wire. All bends were made with the first and second digits to bend the wire around the square beak of the pliers' The resulting specimens were then mounted on cylindrical acrylic blocksll using a domestic adhesivel2 prior to longitudinal reduction on a microprocessor- controlled grinding and polishing machinels illustrated at FIGURE 3 - 7 - The grinding and polishing were carried out on a horizontally rotating disc, with a horizontally rotating specimen holder pressed eccentrically against it. G¡nding pressure, lubricant and amount of water could be manually adjusted' whereas other actions were controlled by the microprocessor. The grinding protocol for specimen reduction is detailed in TABLE 3 - 4.
Abrasive Speed Pressure Lubricant Time srze (rpm) (Newtons) (Min)
320 grit 150 100 water 2 800 grit 150 100 water 1.5 1000 grit 150 100 water 0.5 4000 grit 150 100 water 2
9¡r lap 150 100 Blue 14 3
TABLE 3.4
Grinding and polishing protocol for preparation of wire sPecimens.
,Grit' Note: = Particle concentration of silicon carbide per unit area for'wet and dry' sand-PaPer discs.
11 Araldite D with Hardener HY2960, Ciba-Geigy' Austral¡a Limited' 12 Araldite Epoxy Resin, 5 minute, Selleys Chemical Company, Padstow, NSW, Australia' 13 Abramin, Struers, Copenhagen, Denmark. 14 Blue DP-Nap lubricant, Struers, Copenhagen, Denmark. 59
FIGURE 3 - 7
Abramin (Struers) grinding and polishing machine. 60
When the specimens were approaching a half thickness, progress was checked at xSO magnification, using a reflecting light binocular microscopels.
This was to monitor the removal of scratches caused by the coarser abrasives. As the polish improved, less light was scattered into the eyepiece of the microscope. Once the surface of the wires became non-reflective, it was assumed that a mirror finish had been achieved. This was confirmed by subsequent scanning electron microscopy.
Following successful reduction, the ends of the specimens were carefully annealed by passing through a flame until red hot. This was to avoid changing the microstructure in the vicinity of the bend and allowed the wire ends to be bent in such a way that the specimen could be mounted on an aluminium stub for Scanning Electron Microscope (SEM) examination. Noting the conclusion by Twelftree, (1974) as to the unsuitability of conventional optical metallographic techniques for such investigations, SEM examination was chosen as being appropriate for the investigation of wire microstructure. Each specimen was mounted on an individual SEM stub using a domestic adhesive.l6 This was done, keeping the specimen bend about 3mm above the SEM stub, such that with a variation of stub orientation in the SEM chamber, views of the polished surface and exterior surface could be achieved. However, prior to being etched and mounted on the SEM stub, each specimen was cleaned in an ultrasonic alcohol bath for five minutes, and dried in a blast of warm air, (after Singh, 1993).
To reveal the wire microstructure and characterise the discrete phases of the wire, there was a need to etch each specimen. This allowed for differential
15 Leitz, Wetzler, Germany. 16 Araldite Epoxy Resin,5 minute, Selleys ChemicalCompany, Padstow, NSW, Australia. 61 attack by the etchant to selectively 'unbuild' the structure from the surface downward. The various constituents of multiple-phase alloys, or the sectional planes in pure and anisotropic metals have inherently different rates of solution in the etching reagents. ln alloys, the mechanism of etching is essentially one of an electrochemical nature, due to the difference in potential between the structural components and the etchant; the phase at the higher potential being anodic or electropositive, and therefore tending to go into solution more readily. The cathodic elements remain standing in relief relative to the anodic structures. The anodic phase, having been roughened and pitted, and principally because of shadows when viewed microscopically, tend to appear darkened, (Kehl, 1949).
"By continuing the etching process beyond the time required to delineate clearly the various structural phases present (in a multiphase metal), the component initially electropositive may become ennobled and thus possess with respect to etching all of the characteristics of an electronegative phase' This change in potential, when it occurs, may be attributed to selective deposition of reaction products, or the formation of other kinds of anodic protective layers. In any event, the original cathodic component may be rendered dissoluble, and upon fuilher etching, the structure will become over- etched and poorly defined. By arresting the etching process, however, before the anodic phases become ennobled, electrolytic action in the opposite direction may be avoided with attending better contrast established between the structural phases", (ibid).
,,Etching a pure metal or single phase alloy is a process of chemical solution of the metal by the reagent, where each grain is dissolved at a rate dependent on the orientation of that grain section with respect to the plane of the polished 62 surface. (This develops a) series of well defined facets that are similarly orientated on any one grain section, but which as a group is differently oriented than those on neighbouring grain sections. (With sufficient etch¡ng time) with certain highly active reagents ... these facets will be sufficiently well developed and enlarged to be revealed microscopicatly as individual entities." Etching for a longer time allows for the production of etching pits that are related to the difference in rate along different crystallographic planes, and is a continuation of the development of the etch facet phenomenon. The shape or pattern of etch¡ng pits is associated with the corresponding geometry of the crystallographic system, (¡b¡d). The etch pits are formed at dislocation sites because the strain field surrounding the dislocation causes preferential attack by the etchant. lt is also of note that for metals, etch - pit formation at dislocations is dependent on purity, (Dieter, 1961). lnitially, it was decided to follow previously reported procedures in preparing orthodontic wire specimens for SEM study. Consequently, the wires were etched for seven minutes in a reagent containing 10 per cent concentrated nitric acid , 2 per cent concentrated hydrofluoric acid, and 88 per cent water. The specimens were then washed in water, cleaned in an ultrasonic alcohol bath, and then dried with a blast of warm air, (Singh'1993).
This protocol was found to be unsatisfactory. Despite adequate etching of the external wire surface, there was inadequate resolution of the polished surface microstructure for satisfactory scanning electron microscopy. Subsequent specimens were prepared by washing in an ultrasonic bath of absolute alcohol for 5 minutes prior to being etched. A reagent comprising 1 part concentrated nitric acid, 2 parts concentrated hydrochloric acid, and 2 parts glycerol, (Vander Voort, 1984), was used to etch the specimens in an 63 ultrasonic bath for 5 minutes. The specimens were washed in water and cleaned in an ultrasonic alcohol bath for five minutes prior to being dried in a blast of warm air. This was found to provide a sat¡sfactory etch of the polished surface, (F¡GURE 3 - 8).
A Philips SEM 50517 scanning electron microscope operating at 20 kV, with a Tracor Northern TN 55OO model EDS system microprobe operating with an accelerating voltage oÍ 20 kilovolts (kV) and electron beam current of 3 nanoamperes (nA) was used to analyse and photograph the early specimens. Following flood-water damage to the Centre for Electron Microscopy and Micro-Structural Analysis (CEMMSA), the University of Adelaide, later specimens were studied using the replacement machine, a Philips SEM
XL2O18, operating with an accelerating voltage of 20 kV and capable of a variable beam current of 0 to 200 microamperes (frA). This machine is illustrated at FIGURES 3 - 9 and 3 - 10.
17philips Analytical Electronics, N.V , Eindhoven, The Netherlands. 18Philips Analytical Electronics, N.V , Eindhoven, The Netherlands. 64
- a
:- ¿:=æ
V Spot Magn Dct WD Erp ldl pm kV 6.0 216r SE 34.1 1õ TT-TAT:¿14ÍI :Xscct+ulÏc
FIGURE 3 - 8
Surface of 0.016 inch, (0.4064 mm) 'premium plus' Wilcock wire batch'TT-TAT', following
grinding, Polishing, and etching. 65
>-/'- .t' \¿ /
FIGURE 3 - 9
Philips XL20 Scanning Electron Microscope. 66
\
FIGURE 3 . 1O
Specimen stage of Philips XL20 Scanning Electron Microscope 67
TENSILE TESTING OF WILCOCKTM WIRES
MATERIALS
The Australian Standard, AS 1964 - 1977, classifies resilient orthodontic wires according to six tensile strength ranges, or types. Wires produced and classified by A.J. Wilcock, Australia, include grades that exceed Type 6. These differences are shown in TABLE 3 - 5.
ADSL Tensile range Tensile range AJW classification MPa Tons classification
Type 1 1720 - 2240 (Note 1) Type 2 2241 -2400 145 - 155 Regular
Type 3 2401 - 2550 156 - 165 Regular Plus Type 4 2551 -2700 166 - 175 Special
Type 5 2701 - 2860 176 - 185 SpecialPlus Type 6 2870 + (Note 2) 186 - 195 Premium 3011 - 3166 196 - 205 Premium Plus
3181 + 206 + Supreme
TABLE 3.5
Comparison of orthodontic wire classifications, as classified by the Australian Standard, AS 1964 - 1977, and A.J. Wilcock Scientific and Engineering Equipment (AJW).
Note 1 A.J. Wilcock does not produce wires of this low tensile strength. Note 2 This is currently the highest classification of AS1964 - 1977, lhe measurements for which are based on imperial measurement conversions from the previous Australian Standard T32.
(1 Ton Tensile = 15.444 Megapascals) 68
Samples of wires produced by A.J. Wilcock were submitted to the Australian Dental Standards Laboratory (ADSL) in 1985, for testing in accordance with the Australian Standard, AS1964 - 1977. A confidential report by the Laboratory, details of which are reproduced here by kind permission from Mr A.J. Wilcock, show a discrepancy in tensile strength results achieved by the Laboratory and A.J. Wilcock Scientific and Engineering Equipment. Consequently, the manufacturer decided to retain samples of these batches for future reference. Spools of wire from these batches, still in their original packaging, were subsequently made available for testing as a part of this project. The discrepancies are summarised in TABLE 3 - 6.
Specimen Tensile Strength Details A.J.W A.D.S.L.
Diameter Type / Batch Tons MPa Tons MPa
0.0121" Regular 155.2 2397 150.5 2324 (0.307 mm) }IW-RII
" Reg. Plus 0.01 41 163 2517 148.1 2287 (0.358 mm) DH-RIK
0.0162" Spec. Plus 175.6 2712 167.9 2593 (0.411 mm) V/I-RCO
Prem. Plus 0.016" 196 3027 186.6 2881 (0.406 mm) EH.RKL
Spec. Plus 0.0202" 185 2857 177.5 2740 (0.513 mm) RZ.ROI
TABLE 3 - 6
Tensile strength of Wilcock wires as measured by A.J. Wilcock Scientific and Engineering Equipment (AJW), the Australian Dental Standards Laboratory (ADSL). 69
METHODS
Lengths of wire between 300 and 400 millimetres were cut from the spools provided and mounted in Hounsfield Vee Wire Grips (FIGURE 3 - 1 1) suitable for testing on an lnstron Universal Testing Machinele (FIGURE 3 - 12). The machine consists of two cross-heads mounted in a rigid frame. A load cell, type 2511-104, model A3O-40, was loaded into the upper fixed cross-head, which in turn supports the upper specimen grip. The lower moveable cross- head supports the lower specimen grip, and moves under the control of a programmable screw thread when a tensile load is applied. The load is recorded by the load cell onto a moving chart with a pen recorder. The load cell was calibrated with 10 and 20 kg dead weights, with an accuracy of f0.5 per cent. In practical terms, the load weighing system exhibits no mechanical inertia, and therefore its action does not significantly influence the properties of the sample being measured.
The specimens were tested in accordance with the Australian Standard, AS1964 - 1977. The length of wire between the grips was not less than 50 millimetres, being usually between B0 and 100 millimetres long. A cross- head speed of 0.5 + 0.1 millimetres per minute was employed. lf failure on loading occurred at one of the grips, the test was disregarded and repeated. Six determinations were carried out on each wire type, except for the 0.404 millimetre, 'premium plus', batch number EH-RKL, where there was insufficient length of wire available for more than two tests.
Results were plotted on the pen recorder which had a scale accuracy of 10.25 per cent.
19 lnstronl0 tonne floor modelTT-D, United Kingdom 70
EEP PIN -¡\ t WtRE I
I Aocholt'C i;.: l- -t cd
Vrqr rx vEE cFoov
VEE wrRE cR¡Ps.
60m
FIGURE 3 . 11
Hounsfield Vee Wire GriPs 71
' i' ¡lr.' l
I
- -H ,.=.- .-Tr -Í I- E I . t. - ':rri - != -l\ll
FIGURE 3. 12
Instron Universal Testing Machine 72
ARCH-FORM ANCHORAGE BEND ANALYSIS
MATERIALS ln order to study the effects of anchorage bends in relation to forces in the anterior portion of an arch-wire, some representative arch-form needed selection. Perusal of proprietary catalogues will show a lack of consensus in the orthodontic marketplace as to what represents 'an average arch-form'. Due to the lack of a universal arch-form it was decided to base an arch-form on the method of measurement cited by Rohan, (1982)' and Hazel et al' (1984). For their study, twenty pre-treatment maxillary arch study models were obtained and measured to determine the arithmetic mean for inter-molar width, inter-canine width, and anteroposterior distance from the anterior part of the arch to the inter-canine and inter-molar axes. These measurements allowed the derivation of an 'average arch-form'.
To legitimise this information, measurements were made of twetve maxillary dental casts from material kindly supplied by Professor Grant Townsend of the Anthropology and Genetics Research Group, Faculty of Dentistry, the University of Adelaide. Casts were selected from first year dental students with an intact permanent dentition, (third molars were not considered), a reasonably aligned Class 1 relationship with less than 2 millimetres of crowding, and an absence of cross bites. Those measurements used by Borghesi, (1973) and Rohan (1982) were recorded to derive arithmetic means that were rounded off to the nearest millimetre. One millimetre was added to each measurement to simulate the distance between the buccal surface of the tooth and the arch-wire engaged in a ribbon arch bracket, (atter Rohan, 1982). The results are listed in TABLE 3 - 7. 73
Case No. Sex tc M AP-lC AP- IM
1 Female 36.83 55.72 11.40 31 .10
2 Male 37.17 56.60 9.10 29.43
3 Male 41.03 57.74 14.12 35.30 4 Male 41.59 60.50 12.17 32.81
5 Male 40.01 58.1 4 11.87 34.36
6 Female 40.48 57.32 11.07 32.57
7 Female 37.97 57.68 10.62 28.99 I Female 36.23 56.70 11 67 32.43 o Female 36.32 53.58 7.95 28.79
10 Male 40.41 57.95 10.03 29.98
11 Male 36.05 56.1 0 9.70 30.92 12 Male 41 .19 54.52 10.62 31.73 31.53 MEAN 38.77 5 6.88 10.86 11 32 To nearest millimetre 39 57
12.6 33.1 Borghesi, '73 34.7 51.2 12.8 36.25 Rohan,'82 37.05 54.25
TABLE 3.7
Maxillary cast measurements to determine arch-form dimensions, in millimetres, compared with the results of Borghes¡, (1973) and Rohan, (1982).
Note: lC= distance between buccal surfaces of the canines lM= distance between buccal surfaces of the first molars point axis AP-lC = anterior centre to inter-canine point axis. AP-lM = anter¡or centre to inter-molar 74
Differences between the results in this investigation and those of Borghesi, (1973), and Rohan, (1992), could be due to measurement error, or differences in the samples. Sample differences might contribute to a larger part of the differences seen. The sample in this study was from individuals with a Class 1 occlusion, the other two samples were primarily of Class 2 individuals. lt would seem reasonable to observe broader and shofter maxillary dental patterns in Class 1 individuals. This assumption is in keeping with the noted differences in results.
Based on those measurements in TABLE 3 - 7, a template of the representative arch-form was drawn on a card to the following dimensions:
Arch depth (anterior centre point to inter-molar axis): = 33mm Anterior arch depth (anterior centre point to inter-canine axis): = 12 mm .1 lntermolar width: 58mm 'ìf = 'i! lntercanine width: = 40mm
The representative arch-form is shown at FIGURE 3 - 13.
Three spools of Wilcock2o 0.016 inch (0.4064 millimetres) premium plus wire, (batch 'DH-RKK', manufacture completed on 04 November 1993 with a
measured tensile strength of 3.07 x 103 MPa) were used to fashion arch-forms
based on the template. Wire manipulation was minimised to reduce the effects of cold working and the introduction of excessive strain history into the wire. Where possible all arch-forms were bent with the fingers, pl¡ers being used to form canine hooks and anchorage bends. I I
I
20 A.J. Wilcock Scientific and Engineering Equipment, Whittlesea, Victoria, Australia' t 75
AP.IC 1
AP.IM 2.
rc
IM
FIGURE 3 - 13.
I The representative arch-form rl
'1
lC= distance between buccal surfaces of the canines lM= distance between buccal surfaces of the first molars AP -lC = anterior centre point to inter-canine axis point axis. AP -lM = anterior centre to inter-molar
Notes: 1 Canine hooks 10 mm from anterior centre point. 2 Anchorage bends 10 mm from distal ends.
I
r 76
A variety of hook designs were chosen to examine potential force differences that m¡ght result from such design differences, particularly anterior to the hooks. The different arch forms constructed were:
1 Plain archwires with no intermaxillary hooks or stops,
2 Archwires with intermaxillary'circles',
3 Archwires with intermaxillary 'boot hooks',
4 Archwires with intermaxillary'loop hooks',
5 Archwires with molar stops.
These different arch-wire designs are illustrated in F¡GURE 3 - 14
An experimental apparatus, based on that used by Borghesi, (1973) and Rohan, (1982), was designed such that it was capable of holding a representative arch-wire in two commercially available 0.036 inch (0.914 mm) molar buccal tubes21. A Computer Aided Design (CAD) of the apparatus is shown in FIGURE 3 - 15, and a view of the apparatus holding an activated arch-wire is shown in FIGURE 3 - 16.
The molar tubes, 4.5 millimetres long, and fitted with standard hooks for elastic ('rubber') band attachment, were glued to the brass blocks with a domestic adhesive22. The apparatus consisted of an aluminium 'U' shaped frame, (145mm x 1Oomm) to which were mounted three supports. These were for two adjustable molar tube brass blocks, each able to be rotated for adjustments in molar tube angulation, and adjustable in relation to the distance between them; and, for an incisal pointer that could be adjusted to
21 3M / Unitek Corporation, Monrovia, California, USA. 22 Araldite Epoxy Resin,5 minute, Selleys Chemical Company, Padstow, Australia. 77
FIGURE 3 . 14
Arch-wire designs.
,P' Plain archwire with no intermaxillary hooks or stops
a Archwire with intermaxillary'circles'
B ,. Archwire with intermaxillary'boot hooks'
a ,.
M : Archwire with molar stops FIGURE 3. 15
CAD drawing of the arch-wire testing apparatus by
Mr Ron Jager, Mechanical Engineering workshop,
The University of Adelaide. Width vorioble from 40-60mm
o! C õtolf) >Ool cÐ oç c! oY -l;
o UNIVERSITY OF ADELAIDT Dept. of Mechonrcol Engineering o c ARCHWIRI TEST APPARATUS Scole Full Size DoteT /6/94 P roj' n @ Drown R Joger Sheel File dentol 79
FIGURE 3 - 16
View of apparatus with an activated arch-wire' 80 reflect the arch depth. The molar tube blocks had small tabs at the back to ensure no more than 2.5 millimetres of arch-wire protruded distally from the molar tube and that the wires were completely free to move within the molar tubes. The frame was mounted on a support that was capable of being adjusted up and down by a screw mechanism, and is shown in FIGURE 3 - 17. ln order to test force levels at the anterior portion of the arch-wire, two sharp aluminium knife-edge supports were used. The supports were placed on a levelled electronic scale pan23, that was zeroed, and then used to measure vertical force levels at different points around the arch-wire. Positions for paired knife-edge application were chosen such that they approximated the position of first premolars, canines, lateral incisors and central incisors. The knife-edge supports were glued2¿ in each position for each series of tests to avoid support slippage and validate results between different arch-forms. A single knife-edge was used at the anterior centre of the arch-wire to obtain force values that could be compared with the work of previous investigations, (Borghesi, 1973; sims, 1977; Thornton & Nikolai, 1981; Rohan' 1982; Witlmot, 1983; Hazet et al, 198a; Sampson' 1992)'
23 fx-gOOo Electronic balarrce, A & D Company, Limited, Japan. 24 Araldite Epoxy Resin, S minute, Selleys ChemicalCompany, Padstow, Australia' 81
FIGURE 3.17
Arch-form anchorage analysis apparatus' 82
METHODS
generated This aspect of the study involved an assessment of the initial forces To by a typical 0.016 inch (0.4064 mm) Begg stage 1 intrusive arch-wire' of simulate clinical conditions aS near as practicable, and keep the number variables within manageable levels, certain parameters were kept constant' These were:
'DH-RKK'), a type of wire, (wilcock 0.4064 mm 'premium plus" batch
a arch-form, (as in FIGURE 3 - 13)'
a location of anchorage (tip-back) bend,
a molar tube dimensions,
a molar tube internal diameter,
a molar tube rotation,
a inter-molar tube width,
a heights of the molar tubes relative to each other,
a molar tube height relative to the anterior pointer'
The distance of the molar tubes from each other was maintained at 58 millimetres in order to prevent the incorporation of any expansion or contraction in the arch-wire.
distal Anchorage bends of known degree were placed 10 millimetres from the were ends of each different arch-form. This meant that when the arch-wires in correctly seated in the testing jig, the anchorage bends were 3 millimetres chosen in front of the molar buccal tubes. The anchorage bend angles were gain a order to conform aS closely as possible to previous work and thus 83 measure of comparison. Anchorage bend angles used by previous authors are given in TABLE 3 - 8.
Author Year Bend angles used (degrees) Borghesi 1973 10, 20, 25, 30, 40, 45, 50 Sims 1977 30, 45, 60,80
Thornton & Nikolai 1 981 30,45,60
Rohan 1 982 20, 30, 40, 50, 60
Willmot 1 983 10, 20, 30, 40, 50, 60
Hazel et al 1 984 45
Sampson 1 992 20, 40, 60
TABLE 3.8
Degrees of anchorage bend investigated by previous authors.
The angles used for this investigation were 15, 30, 45, and 60 degrees' These angles were considered to resemble not only prev¡ous work but also to be clinically relevant in so far as it was considered questionable that many clinicians can 'judge by eye' smaller angles with any degree of accuracy or consistency.
Anchorage bends were chosen to allow as much correlation as possible with other studies. This was despite claims by Lew, (1990)' that anchorage curves are more efficient in attaining incisor intrusion and preventing mo¡ar extrusion. This was explained by suggesting that the effect of an anchorage 84
curve is expressed along the entire length of the arch-wire compared with the expression of intrusive ability of an anchorage bend at a single point'
Four arch-wires were made of each type, such that there was one for each
degree of anchorage bend per arch-form.
Following placement of the arch-wires into the testing jig, the anchorage bends were activated by winding the screw mechanism of the jig. This meant that the arch-wire was 'wound down' onto the knife-edge suppot'ts. The arch' wires were elevated until the anterior centre point of each arch-wire was level with the incisal pointer, simulating arch-wire engagement into anterior Begg brackets. Strain on the arch-wires could be adjusted by altering the relative height of the molar tubes. Reproducibility of the degree of force applied by the jig was found to be very accurate because of the precise machining of the t'l ,r steel pointer and the smooth calibrated action of the screw mechanism.
The amount of force, or load imposed on the knife-edge supports, was directly related to the stress delivered by the wire. since this force was directly measurable, the stress delivered by the wire was, therefore, able to be determined. This was the conceptual basis of the apparatus.
This investigation ailowed comparison of load / deflection characteristics of arch-wires with different inter-maxillary hook configurations and varying degrees of anchorage bend.
I t 85
CHAPTER 4
RESULTS AND DISCUSSION
THE SEM EXAMINATION OF WIRES
WIRE MICROSTRUCTURE
During the fabrication of wires, the cast metal or ingot is rolled and drawn down to size in stages. Between each stage the wire is heated to the re- crystallisation temperature of the metal, thus producing smaller, equiaxed grains. The drawing process elongates these grains into long fibrous structures, the long axes of which run in the same direction as the draw' This ü tii elongation is known as cold working (by plastic deformation), and the finished product is known as a wrought structure, (Kohl, 1964). The directionality in mechanical properties produced by rolling and drawing can have an important effect on wire forming by the clinician. The alignment of the material in preferred directions produces a fibrous structure that is characterised by a 'texture', (Kusy & Dilley, 1984). Mechanical texture has little effect on formability, however, crystallographic fibering, or preferred o¡entation of the wires' crystal lattice, can have a large effect on formability. For sheet metals, bending across the fibre orientation is mechanically simpler than bending parallel to the crystal orientation, (Dieter, 1961), whereas this problem is unlikely to occur with wires, since, in practical terms, wires are
always bent across the fibre orientation.
I
r 86
Extra-hard drawn wires are generally used for fixed orthodontic treatment. These wires have undergone a great amount of work hardening during the drawing process, with an attendant distortion of the microstructure. Simple microscopic examination of wires that have undergone basic etch¡ng cannot distinguish discrete phases. The visualised patterns are often confounded by rather empirical heat treatments that add further complexities to the wire's appearance. tf an etching process could distinguish discrete phases and characteristics of the wire, they could be correlated with clinical performance, and success or failure of wires be attributable to their actual microstructure. Clinician's requirements might then be evaluated and fulfilled, (Williams & Fraunhofer, 1971). lt has been argued that simple etching procedures, designed to dissolve certain phases selectively from stainless steel wires, require considerable experience and skill, and are not easily applied to complex structures. As an alternative, a more sophisticated method of electrochemical etching, using a potentiostat, is recommended, (ibid)'
Electrostatic etching with the potentiostat requires each specimen to be mounted in acrylic during the etching process. The requirements of specimen manipulation during scanning Electron Microscopy precluded mounting specimens in acrylic blocks, and it was decided to use the conventional metallographic etching regimes after Vander Voort' (1984)' These 'conventional' methods were used, cognizant of the comments of Witliams & Fraunhofer, (1971), regarding the ability of such methods to
discern different phases in the etched specimens.
was considered possible to I By an examination of bent specimens of wire it I These I observe any work hardening effects subsequent to manufacture. additional effects may provide information about the reasons for occaslonal
I 87 faiture of the wire during clinical manipulation. Additionally' such investigation will provide information as to the early microstructural changes that occur during the drawing process itself, (Williams, 19741.
It has been suggested that no critical examination of wires is complete without an analysis of the composition of the wire, (Williams, 1974), since the nominal composition may vary from that as specified by the manufacturer. An analysis of batch TT-TAT of Wilcock2s 0.016 inch, (0.4064 millimetre) ,premium plus' was carried out using the X-ray micro-analysis capability of the SEM. The results, using the 'ZAF Quantification Method' from the 'EDAX' software26, are shown ]n TABLE 4' 1.
Element K ratio Weight % Atomic %
Chromium 0.1974 17.120 18.235 lron 0.7232 74.090 73.473 Nickel 0.0795 8.790 8.291
TABLE 4 - 1
X-ray micro-analysis of electron 'K' shell energy to
determine relative concentration of elements in
A.J. Wilcock batch TT-TAT 0.016 inch (0.4064 mm) 'premium plus'.
2s A.J. Wilcock Scientific and Engineering, Whittlesea, Victoria, Australia. 26 EDAX lnternational, Mahwah, New Jersey, USA' 88
Label: Wire TT-TAT (bulk) heset: None Live Timc :40 l5:28:48 8-24-94
eK
1.00 2.00 3.00 4.00 5.00 6.00 7.00 8.00 9.00 FS : 563 CPS: 1030 Cnts :49 KeV:0.53
FIGURE 4 - 1
X-ray micro-analYsis of the
main bulk of wire in batch TT-TAT. 89
The results in TABLE 4 - 1 indicate a fairly typical '18-8' stainless steel. lt should be noted that no attempt was made to determine the presence and concentration of minor trace elements due to the difficulties in separating these from significant background 'noise' or backscatter in the analysis and aluminium contamination from the SEM stub. The micro-analysis is in graphical form at FIGURE 4 - 1.
,,One of the most convenient ways of representing the effect of various elements on the basic structure of chromium - nickel stainless steels is the Schaeffler diagram. lt plots the compositional limits at room temperature of austenite, ferrite, and maftensite, in terms of chromium and nickel equivalents. At its simplest form, the diagram shows the regions of existence of the three phases for iron - chromium - nickel alloys. However, the diagram becomes of much wider application when the equivalents of chromium and of nickel are used for the other alloying elements. The chromium equivalent has been empirically determined using the most common ferrite-forming elements:
Cr equivalent = (Cr) + 2(Si) + 1.s(Mo) + 5(V) + 5.5(Al) +1.75(Nb) + 1.S(Ti) + 0.75(W)
Similarly, the nickel equivalent has been determined from the familiar austenite-forming elements :
Ni equivalent = (Ni) + (Co) + 0.5(Mn) + 0.3(Cu) + 25(N) + 30(C)
All concentrations are expressed in weight percentages. The large influence of carbon and nitrogen are of particular note", (Honeycombe, 1981). A Schaeffler diagram is shown at FIGURE 4 - 2. By using the values obtained by X-ray micro-analysis of the wires under consideration, as shown in TABLE 4 - 1, it would seem that the wire could be a combination of austenite, 90
25 Austen ite
20 s Austenile Austenile + c + q) martensrle ð- ferrile ñ 15 .¿ q o) -q)x z.9 10 Martensite Austenite + marlensrle 5 + ô- ferrite Ma rtensrte ð ferrite + - -f err 0 0 5 10 15 20 25 30 35 40 Chromium equivalent (%)
FIGURE 4 - 2
Schaeffler diagram showing effect of alloying elements on the basic structure of chromium - nickel steels, (from Honeycombe, 1981).
'X' : Composition of Wilcock'special plus' as determined by Newman, (lvou, 91
martens¡te, and ô-ferrite. Ferrite is referred to as the 'ô' phase, since, in chromium - nickel stainless steels, ferrite can have a continuous existence at room temperature, (Honeycombe, 1981). However, without a full elemental analysis, the alloying effects of those elements in low concentration, such as carbon or nitrogen, cannot be considered. lt should be noted that the capabilities of the available X-ray micro-analysis2z, only allow recognition of elements with an atomic weight greater than or equal to oxygen' Therefore, no allowance for the effects of carbon or nitrogen can be accommodated. Given the wire is an Alsl Type 302, (Wilcock, 1994), it is likely that the chromium and nickel equivalents play an important role in determining the phase composition of these wires.
Newman, (1963) presented an elemental analysis for "Australian wire". Given the date of publication, it is likely that the wire was of a lower tensile grade, perhaps 'special' or'special plus'. While the tensile strength of these wires is significantly less than for 'premium plus', (¡e a tensile strength range of 2.5 - 2.86 x 103 MPa versus a strength of 3 x 103 MPa), the elemental composition is unlikely to vary significantly. Using the f¡gures of Newman' (1963), shown in TABLE 4 - 2, and substituting these into the 'equations' for chromium and nickel equivalents, the following figures are extant: cr equivalent = (cr- 18.21) + 2(Si - 0.48) + 1.S(Mo -0.21) + 5(V - 0) +5.5(Al -0) +1.75(Nb-0) +1.5(T¡ -0) +0.75(W-0)
Ni equivalent = (N¡ -8.97) + (Co-0) +0.5(Mn -1.17) +0'3(Cu-0) + 25(N - 0) + 30(C - 0.0e5)
1e: Cr equivalent = 19.485 weight Per cent.
Ni equivalent 12.405 weight Per cent.
27 EDAX lnternational, Mahwah, New Jersey, USA. 92
Element Wilcock wire, (Newman, 1963)
Carbon, (maximum) 0.095 Manganese, (maximum) 1.17 Phosphorus, (maximum) 0.022 Sulphur, (maximum) 0.032 Silicon, (maximum) 0.48 Molybdenum 0.21 Nitrogen, (maximum) Chromium 18.21 Nickel 8.97
TABLE 4 - 2
Compositional analysis of 'Australian' wire by Newman; (1963).
The point at which this composition lies on the Schaeffler diagram, (FIGURE 4 - 2) is indicated by 'X'. Accordingly, it is suggested that the wire, as tested by Newman (1963) is principally austenite, but is a duplex sta¡nless steel since it comprises about 5 per cent ô-ferrite. The relative percentages of ô-ferrite in a Schaeffler diagram are shown in the text by Colombier & Hochmann' (1967). This finding of õ-ferrite in Wilcock wire is in agreement with wiltiams, (1974), who noted about 16 percent õ-ferrite, and suggested that its presence bore no correlation with observed mechanical properties' wiiliams, (1g74) has indicated that high tensile wires produced by A.J. Wilcock of Australia exceed the bounds of a pure austenitic material. They are often ferromagnetic, due either to the presence of ferrite and / or 93
martens¡te, or, due to the effects of preferred crystal orientation that ls a result of the drawing process. He showed that other elements have a chromium and nickel equivalent effect and therefore alter the relationship and amounts of constituent phases present. Some elements will have a direct effect on the phases to be found, for example molybdenum is a known ferrite-former in stainless steels. ln addition, he suggested that Wilcock wires "have an austenitic base stock, (and that) work hardening of the base material, in the process of drawing down, clearly produces changes that were recorded by an increase in ferromagnetic susceptibility". This identified ferromagnetic feature of heavily drawn AlSl Type 302 austenitic wires is in agreement with comments made by Witcock, (1994) and may be due to deformation- induced martensite, formed from metastable austenite that is the result of extensive cold working, (Khier et â1, 1988), or, the presence of low concentrations of ô-ferrite, Colombier & Hochmann, (1967).
Williams, (1974\ cites Sokolov et ât, (1966) who suggest that the extensive cold working of high tensile wires leads to the formation of ô-ferrite rather than martensite. The formation of stable maftensite in austenite is augmented by the presence of most alloying elements, particularly carbon and nitrogen, but importantly, is inhibited by stress, (Honeycombe' 1981)' such as occurs with cold extensive cold working during the manufacture of Wilcock wires.
Colombier & Hochmann, (1967), point out that some disagreement exists as to the calculation of chromium and nickel equivalents. They suggest that the presence of õ-ferrite can only be confirmed by microscopy, and that for the composition of Wilcock wire, as determined by Newman, (1963), some authors would consider the wire to be austenitic, devoid of õ-ferrite. 94
BENDING EFFECTS FROM DTFFERENT TYPES OF PLIERS lnitially three different types of pliers were used in an assessment of the effects on bending Wilcock 0.016 inch (0.4064 mm) 'premium plus' wire. After the commencement of this project, information from the manufacturer, (Wilcock' 1gg4), led to the decision to compare the results of the initial three pliers28,zg,3o against the performance of pliers that had been developed by Mollenhauer of Melbourne, Australia. These Plierssl have been designed with tungsten carbide inserts in the beaks and are claimed to reduce the incidence of high tensile wire fracture. The inserts are flat, with edges that have been rounded and polished with an impregnated rubber wheel. Due to their appearance, these pliers have been nicknamed 'dolphin' pliers by the manufacturer. Given the results that had been observed with the initial three plier types, fufther samples of wire were fabricated with 40 and 60 degree bends for testing with the 'dolphin' pliers.
The wire was initially viewed as an un-prepared specimen. This gave an insight into some of the surface irregularities in the 'as received' condition in the vicinity of a bend of about 40 degrees. The wire was bent with stainless steel tipped, light wire pliers, and an example is shown at FIGURE 4 - 3. The main feature to note in this specimen is the overall roughness of the wire's surface. There are areas of scale visible, and the longitudinal drawing lines from the die are evident. A small cylindrical body, about 5 x 20 pm, and of unknown origin, is marked 'U'.
28 Catalogue number5162, Crovana, Germany. 29 Catatogue number 442, Rocky Mountain Orthodontics, Denver, Colorado, USA' 30 Catalogue number 105, Dentronix, lvyland, Pennsylvania, USA'
Whittlesea, Victoria, Australia. 95
FIGURE 4. 3
Surface irregularities in the vicinity of a 40 degree bend in 'as received' TT-TAT,0.016 inch, (0.4064 mm) 'premium plus"
'U' : Cylindrical body, 5 x 20 pm, of unknown origin' 96
The prepared specimens that were bent w¡th the stainless steel light wire, or '139'pliers showed little or no unusual characteristics when viewed by scanning electron microscopy. Both the internal and external wire surfaces in the vicinity of the bend showed a smooth transition around the arc from one straight section to the next. A typical example of what was seen is shown in FTGURE 4 - 4. Of note is the increased pitting seen along the 'core' of the specimen. This longitudinal band of pitting is 25 to 35 pm wide. lt is likely to have similar micro-hardness properties to the 40 pm core as described by Williams, (1 974). This 'core' feature is likely to be due to strain inhomogeneities, probably as a result of redundant deformation that occurs during drawing. The redundant deformation can be considered as that required to change the shape of the wire during deformation, but not reflected in the final wire size. The amount of redundant work performed on the wire during drawing is dependent on the approach angle in the die and on the reduction in area, and, to some smaller extent, on the initial wire diameter, and on the extent of prior deformation, (Davies, 1987).
Redundant deformation has a number of effects that are detrimental to wlre properties. There is an increased risk of internal central bursting during drawing, since the drawing stresses become increased, and the total residual stress and strain in the wire are increased. Variations in propefties across the wire cross section occur, (¡b¡d). This may give rise to an increased failure rate of the wire, but in the least instance, helps to explain the 'greenstick' failure pattern seen in Wilcock wire, particularly when it is manipulated in contradiction to the manufacturer's specifications, (Wilcock, 1988, 1989). 97
I ï
I \ I lrr I \ t r I ;l 'T1 | , '' ¡' I ì !' I ¡ I ll I i i lt I I { i
, I I ,,1 I ¡ I I 100 pm c(, V Spol Magn Det WD F xp CkV60 30Ox sÊ 338 I SS/I.W +4-Xsecl
FIGURE 4.4
Specimen showing increased pitting along a longitudinal 'core'. 98
FIGURE 4 . 5
Specimen displaying reduced etch¡ng effect. 99
2'
t-;
pm V Spot Magn Dct WD Erp 1ü) kV 6.0 fþ¡ SE r23 12
FIGURE 4 - 6
Specimen displaying increased etching effect
Note the persistence of the increased etching in the longitudinal 'core' area. 100
The 'core'feature was not found in all specimens, and it is assumed that small variations in etching technique or susceptibility to etching are responsible for different degrees of pitting seen. This difference is shown by comparing the etch pattern in FIGURE 4 - 4 with those patterns seen in FIGURE 4 - 5, (lesser etching effect) and FIGURE 4 - 6 (greater etching effect). However, the specimens in FTGURES 4 - 5 and 4 - 6 were etched simultaneously. As such, a factor other than time must have been responsible for the different effects seen. Such factors might include some burnishing of the 'lesser etched' specimen during polishing or, differences in the structure, (and thus susceptibility to etching), of the wire along the batch length. tn one sample of the specimens that were bent with the stainless steel light wire pliers, a distortion was found on the internal surface of the bend. There appeared to be some flattening of the internal surface of the wire at the maximum inflection of the bend. Next to this was a more pronounced part of the bend before the wire straightened distal to the bend. The appearance of this distortion is commensurate with excess plier holding pressure during the bending process and is illustrated at FIGURE 4 - 7.
The distortion shown in FIGURE 4 - 7 suggests that the sharpest part of the bend is where the edge of the rectangular plier beak was positioned during bending, ('A'). The small size of the distortion, compared to others to be shown later, suggests that one of two reasons may be extant. Firstly, the amount of plier holding pressure may have been small, or, secondly, the relative hardness of the plier beak may have only been sufficient to cause the slight distortion seen. A combination of the two factors may be valid. However, given that only one specimen in this series was seen to have this type of edge defect, excessive plier holding pressure is the most likely cause. -_ _- L, t -
V Spot Magn Det WD Erp õ0 Um ì20.0 kV 6.0 Sfi)x SE 16.A 2 pllcr ofrect
FIGURE 4 - 7
Stainless steel light wire plier effect on concave surface of a 50 degree bend.
'A' : Point of plier beak applicat¡on 'l02
The next series of wire specimens were bent us¡ng a pair of 'Tweed' torquing pliers, commonly referred to as '442' pliers. These pliers feature two flat stainless steel beaks. As with the stainless steel tight wire '139' pliers, all but one specimen displayed a regular transition of the wire form around both the convex, external surface and concave internal surface of each bend. The single specimen that showed the edge defect was a 30 degree bend, less likely than a larger angled bend to cause undue strain on the internal surface. Due to the isolated occurrence of this phenomenon, excess plier holding pressure is considered as the most likely ætiology.
This individual specimen is detailed in FIGURE 4 - 8. As a more highly
etched specimen, the longitudinal lamellar structure of the wire is evident. ln the body of the wire adjacent to the surface plier defect, there appears to be a smooth and graduat transition of the longitudinal structure. The 'longitudinal lines' of etched pits, initially run parallel with the surface. Moving away from the surface, a gradual change occurs, until the lines follow the general curve of the wire about 20 pm from the surface. This gradual smooth transition is
suggestive of a lack of additional heavy plastic compressive strain in the area, but does suggest some compression of the local structure. These edge defects, although small, could probably be avoided if the plier holding pressure had been lighter in these specimens. This complies with the advice given by Mollenhauer, (1990).
Using a potentiostat with an austenite etching reagent of sulphuric acid and ammonium cyano-sulphide, Williams, (1 9741 showed that where compressive strain of austenite had occurred during bending, certain features were found in the microstructure. The formation of ô{errite as a result of cold
wo s accompan v 103
F¡GURE 4. 8
Example of plier edge defect in 30 degree specimen
bent with Tweed torquing Pliers' 104 austenite. The grains on the outside of the bend do not suffer any apparent change, whilst etching of those crystals on the inside or compression side of the bend removes phases to display intra-granular slip planes upon which precipitates have formed and have been removed by the etching reagent". These areas of austenite - ferrite transformation often show Widmanstätten side plates that form perpendicular to the grain boundary. These side plates are classically found in cold worked structures. "Where the compressive strain is great, (such as in fracture situations), the crystals become oriented so that the slip planes become aligned at right angles to the direction of the compressive force". Generally however, there was great difficulty "in recognising individual grains in the photomicrography of high tensile stainless steel wires, and it was not possible to visually demonstrate the differences between the outside and inside of the wires, (¡b¡d).
The examination of specimens bent using the tungsten carbide tipped light wire '105' pliers showed a more frequent, and more severe plier edge defect. This finding was not restricted to those specimens with any particular degree of bending. Examples are shown in FIGURES 4 - 9 and 4 - 10.
As well as the increased severity of the plier edge defect, there is an increased distance into the body of the wire prior to the longitudinal pattern being re-established. While this distance for stainless steel tipped pliers, ('139' and'442'), was of the order of 2O pm, in the tungsten carbide tipped light wire plier specimens, the distance was generally 40 to 60 pm, or two to three times as much. Although it could be conjectured that this may be the result of excess plier holding pressure during bending, extreme care was taken during specimen preparat¡on to avoid this very problem. The most r:r-^t., at *Jra innraaoa in hnth carraritrr anrl fra¡,ttanôv nf nliAf eflOe trr\vu vLrsvv - -'-.J defects is due to the hardness and keen edge found on these types of pliers. 105
,i,åæ È--ra
Acc V Spot Magn Det WD Exp 50 ym æ0kV60 500x sE 350 1 TT-TAT 105'50deg:2
FIGURE 4 - 9
Example of plier edge defect in 50 degree specimen
bent with tungsten carb¡de tipped light wire pl¡ers. 106
-ì¡*
L=-
-__.-_t _ Ë------:-:-
Acc V Spot lvlagn Def W[r E xp [ ]] iro pnrr 20ûk\r60 1!000x st '13 !l ,¡ lf I lllll 10h ,40t1eg #4
FIGURE 4.10
Plier edge defect as viewed from the side of an ¡ncompletely polished specimen prepared with tungsten carbide tipped light wire pliers. 107
The other feature to note with these specimens bent with the tungsten carbide tipped light wire pliers was the increased number of inclusions. This increased frequency is probably due to in increased density of inclusions in this part of the wire's length. lnclusions will be discussed more fully in the following section, 'Defects and inclusions'.
Given the increased frequency of plier edge defects in specimens prepared with tungsten carbide tipped light wire pliers, it was decided to compare these results with specimens prepared with the tdolphin' pliers that feature tungsten carbide tips with carefully rounded edges and polished tips. Due to the smoothness of the tips, gripping and manipulation of the wire is difficult without some degree of practice. This meant that there was a conscious increase in plier holding pressure during specimen preparation compared to the other pliers used. Despite this increased pressure, there was no evidence of plier edge defects in any specimens prepared this way. A typical example is shown in FIGURE 4 - 11.
It would seem reasonable to make two suggestions regarding the bending of high tensile Wilcock wires. Firstly, to avoid excess plier holding pressure, and, secondly, to avoid the use of tungsten carbide tipped pliers unless they are well polished and have rounded edges. These findings are in accord with the suggest¡ons of the manufacturer Wilcock, (1988' 1989, 1993)' and of Mollenhauer, (1990). 108
FIGURE 4.11
Typical example of specimen prepared with 'dolphin' tungsten carbide tipped pliers.
Note the residue on the polished surface from incomplete washing in alcohol after etching. 109
DEFECTS AND INCLUSIONS
Internal defects in wire include cracks due to seams or pipe in the hot-rolled ingot-derived starting material, or due to a feature known as cupping. Cupping is the rupture of the centre, or core of a wire when subjected to a tensile force, such as during drawing. Cupping is recognisable either by a localised necking during the drawing process, or by a characteristic 'cup and cone' fracture pattern when the wire is broken. These fracture patterns are often associated with wires formed with high angle dies and high friction, (Dieter, i961). No fractures of this type have been evident in observations of specimens of Wilcock'premium plus' wire under investigation.
"surface checking can occur due to poor lubrication, while longitudinal scratches are caused by a scored die, by improper surface lubrication, or by abrasive parlicles being drawn into the die with the wire. An example of this surface scoring is at FIGURE 4 - 12. Slivers and seams result from cold shuts and blowholes in the hot-rolled starting material. Surface discolouration and ground-in oxide result from improper cleaning of the hot rolled bar or rod," (ibid).
Harcourt & Munns, (1967) suggest that surface defects on wires are the most likely cause of fracture and could be introduced during manipulation. They concluded that the pliers used should be of the "correct form" and shape and discarded when worn or damaged. lt is likely that these authors are commenting on surface defects larger than those observed in FIGURE 4 - 12, since low magnification light microscopy was used to study wires of lower tensile strength, such as those utilised in the fabrication of removable appliance components. 110
c V Spot Magn Det WD Erp 50 pm 0hV60 ¿100x sE 26.5 6 ss/Lw60d #+xt.drawin efrect
FIGURE 4 - 12
Longitudinal surface markings from the wire drawing process of 0.016" (0.4064 mm) Wilcock'premium plus', batch TT-TAT 111
Most structures have built in flaws, either mechanical, such as re-entrant angles, or metallurgical, such as inclusions. These flaws act as a point of concentration for stresses and strains, often leading to cracking. Premature fracture at stresses and strains below the nominal yield strength may occur, ie there is an increased brittleness at the flaw locus, (Colangelo & Heiser'
197 4). lnclusions are originally associated with a metal in the molten state and subsequently appear in the casting or ingot as one of two distinct types; exogenous or indigenous. Exogenous inclusions result from the accidental entrapment of foreign bodies, such as refractory particles from the furnace lining, ladle and moulds, and are usually large but few. lndigenous inclusions occur naturally in metals, especially steels, and result from changes in temperature or composition during normal steel making practices, such as aluminium deoxidation or silicon deoxidation. These inclusions are by definition non-metallic, and usually possess characteristics markedly different from the bulk material and can be considered as an aggregate / matrix system. Consequently, a number of factors will have a bearing on the metal's performance. These factors include the volume percentage, shape, orientation, and mechanical properties of the ¡nclusions, and the d¡rection of principle stresses with respect to this orientation. Colangelo & Heiser' (1974) suggest examples of inclusions in steel making such as:
a manganese sulphide, a iron sulphide, o manganese silicon trioxide, a manganese oxide, a iron oxide, a silicon dioxide, 112
FIGURE 4 - 13
Photomicrograph of Al - Ti - O inclusion identified in batch TT-TAT. 113
An example of an inclusion is shown at FIGURE 4 - 13. X-ray micro-analysis of electron 'K' shell energy levels identified the inclusions as comprising aluminium, titanium and oxygen. Microanalysis of the atomic composition of the inclusion is at FIGURE 4 - 14. lt is very unlikely that the inclusions are artefacts of the polishing process, since the inclusions (2 pm) are much smaller than the finest particles of the industrial diamond polishing medium (9 pm), of an irregular pattern, and have a distinctive X-ray energy signature as described. Examples of polishing particles are shown in FIGURE 4 - 15.
Wiiliams, (1974) showed the presence of inclusions in Wilcock wire and noted their tendency to be aligned in longitudinal 'strings' parallel to the long axis of the wire. This pattern is commensurate with the heavy cold working during the drawing process.
Singh, (1993) also showed the presence of inclusions associated with pitting in various nickel-titanium wires, and described them as "precipitates of titanium dioxide".
Assessment of the properties of inclusions, such as micro-hardness testing, reveals that many types of inclusion are refractory in nature and tend to be hard and brittle. However, the sulphides are relatively plastic and deform readily. Consequently the relative percentage volume of inclusions is modified by other factors, such as the type of inclusion, when considering the overall mechanical properties of a specimen. The relative level of inclusions present in any specimen needs determination prior to any assessment of effect on properties. This is done by several methods, all of which look at the number of inclusions present on a representative surface of a polished samp cr 114
Label : Wire TT-TAT (inolusion)
Preset: None Live Time :40 15:26:04 8-24-94
cK
K
1.00 2.00 3.00 4.00 5.00 6.00 7.ffi 8.00 9.00 FS :289 CPS :830 Cuts :94 KeV:0.53
F¡GURE 4.14
X-ray m¡cro-analysis of the inclusion found in batch TT-TAT. 115
-l lâ :'...::: ,-... ì:--- -l¡ I
.Ç€, .a
t.*..'-. j.+1 k'.- 'a .1 - ;5-- ìtP 20 pm Spot Magn Det WD Exp V 442:30deg:*l parts kV 6.0 1000r sE 11.2 6 F¡GURE 4 - 15 Example of polishing Part¡cles. 116 relative percentage of inclusions. lnclusions are categorised as either sulphides, aluminates, siticates, or oxides, (Colangelo & Heiser, 1974). Given the X-ray micro-analysis of the current specimens, (FIGURE 4 - 14), the inclusions found would best be described as oxides, since the aluminium stubs used to mount the specimens are likely to have contributed to significant 'noise' levels of aluminium in the analysis. The identification of inclusions has been noted previously by Harcourt & Munns, (1967). However they believed that non-metallic inclusions did not detract from the mechanical properties of the wire, particularly if they did not occur in significant quantity, and, where the wire, in the form of a helical spring, had been machine formed. The influence of inclusions on mechanical properties has been broadly reviewed by Thornton' (in Cotangelo & Heiser, 1974) who states that 'Îhe effect of inclusion content on strength is not clearly defined. The ultimate tensile strength and yield strength appear to be relatively unaffected by inclusion content over a wide range when tested in air, but they do show some degradation when tested in a corrosive medium". This is supporled by the work of Toms, (1988) who showed that certain mechanical properties of both stainless steel and nickel- titanium orthodontic wires were adversely affected by exposure to a simulated orat environment over time. The simulated environment included variations of pH, (from 3 to 7), and showed a reduction in ultimate tensile strength, Young's elastic modulus, and 0.2 per cent yield strength. Williams, (1974) suggests that martensite or õ-ferrite phases of stainless steel may be responsible for corrosive properties, and are thus undesirabte constituents of materials for intra-oral use. To this end, the stable AtSl Type 302 austenitic stainless steel, (known as 18:8) is considered suitable. 117 Although fatigue life for a material is not well defined, since it is a statistical value for a given specimen, there is some evidence to suggest a possible link between inctusions and metal fatigue. Fractures, as evidenced by macroscopic crack growth, have been observed to be the sum of two features: firstly, the advance of a crack front through the matrix, and secondly, the advance of a crack front through and around constituent particles or inclusions. The effectiveness of a particular inclusion acting as a locus for stress concentration with a subsequent local¡sed fracture depends on several features: o The distribution of inclusions can have an effect. For example, those spaced close together or in a line, can act aS a single unit, thereby raising the local stresses very efficiently. a Inclusion shape and orientation. For example, elongated .I ru tar ',þ inclusions that are aligned parallel to the principle stresses are less damaging than those that lie normal to the principle stresses. Similarly, rounded inclusions are less damaging than angular ones. a The physical characteristics of the inclusions. For example, hard refractory inclusions such as the oxides are more damaging than the more ductile sulphide inclusions, (Colangelo & Heiser, 197 4). It should be noted that of all the specimens studied in this investigation, only a limited number of such inclusions were found. This may be for one of a number of reasons. One possibility is that this batch (TT-TAT) of 0.016 inch (0.4064 millimetre) Wilcock 'premium plus' wire is particularly free from inclusions. More likely however, is that there was a loss of inclusions by me u etched in an ultrasonic bath for five minutes and cleaned in an ultrasonic r 118 alcohol bath for a further five minutes both before and after etching, this is considered the likely possibility. This possibility is supported by the observation that the polished surface of all specimens contained varying degrees of pitting as viewed under SEM. These pits can be partly explained by the etching process itself. Preferential dissolution occurs in areas of higher ductility, or, variation in grain orientation for single phase alloys. For multi-phase alloys, preferential dissolution occurs in electropositive areas. These etching effects have been described in Chapter 3, Preparation of wires for SEM. However, some of the pits might also be explained as the I remaining niches that contained inctusions prior to ultrasonic etching and cleaning, and subsequent SEM analysis. Determining the difference is not I possible without the presence of inclusions in any particular pit. fl Itj il J FRACTURES There are few repofts of fractures of 'Wilcock'wire (also known as 'Australian' Wire32) to be found in the literature, (Lee, 1992), however, it has been suggested that although infrequent, failure by fracture during manipulation or clinical distortion can be associated with a particular batch of wire, or length of wire on a spool, (Williams, 19741. Fractures Seem to occur more frequently during the forming of right angled bends, (¡b¡d), and may be related to the type of plier beak used during forming operations. Tungsten carbide tipped pliers are not recommended, Wilcock, (1988). I 32 Manufactured by A.J. Wilcock, Australia, as ma¡teted by TP Orthodontics lnc., La Porte, lndiana, USA. 3 119 Wilcock, (1988; 1989), has suggested that high tensile wires should be bent around the flat beak of light wire pliers, such that a moment arm is created between the point at which the wire is held between the beaks of the pliers and the 'bending fingers', using the edge of the flat beak as a fulcrum. ln addition, he suggests that the square beak edge should be slightly rounded to reduce the stress concentration occurring at the bend in the wire. Finally, he suggests that lightly warming the wire to about 40 degrees by running the wire through the fingers prior to bending corresponds with the ductile-brittle transition temperature and therefore reduces the likelihood of fracture. It should be noted however, that austenitic stainless steels normally do not have a ductile-brittle transition temperature, (Honeycombe, 1981), at or near room temperature. Some other explanation for a reduced rate of failure with 'pre-warmed'wire must be extant. Another suggestion has been to alter the shape and form of the pliers used to bend high tensile wires. The use of pliers with two flat polished beaks and rounded edges, such that there is no point contact with the wire, as in a round beak, has been suggested by Mollenhauer, (1990). ln addition, the wire should be held gently with the pliers, and bent slowly with the fingers that hold the wire about 20 millimetres from the pliers, (¡b¡d). Fracture type is often described in relation to the deformation that occurs immediately prior to failure, ie brittle or ductile; the crystallographic nature of failure, ie shear or cleavage; or the appearance, ie fibrous or granular. The actual mechanisms leading to failure are little understood, although an I important factor in fracture ætiology is the stress distribution, (Rohan, 1982). '|.20 It was found during specimen preparation that the wire being studied showed no evidence of breakage. Thus, in order to study fractured specimens, deliberate fractures were made by holding the wire firmly with tungsten carbide tipped light wire pliers and bending the wire very quickly with a snapping motion. This produced forces that are significantly in excess of those used clinically. FIGURE 4 - 16 shows the typical results of a deliberate fracture of 0.016" (0.4064 mm) 'premium plus' wire. For ease of description, the point at which the failure occurs on the convex surface will be termed the proximal end of the fracture, and the distal end of the fracture will be where it tapers into the body of the wire. !nspection of the fracture site shows an initial cross-sectional failure at the proximal end that reverses direction at the wire's mid point. The distal fragment shows a flattened area just distal to the initial failure point, representative of elongation. The proximal fragment has the appearance of â 'Y', the arms of which are the cross-sectional parts of the fracture, and the foot of which is the longitudinal part of the fracture. The proximal portion of wire has superimposed on the cross-sectional failure, a longitudinal fracture pattern, shown in greater detail in FIGURES 4 - 17 and 4 - 18. Where the fracture patterns join there is a transverse narrowing of the wire, indicative of transverse stresses. Distally, the remainder shows a generally longitudinal 'peeling' pattern as is occasionally observed clinically. It would seem reasonable that the type of strain inherent in the wire should match its failure pattern. To this end, it is considered that longitudinal The I stresses are tensile at the surface and cause cross-sectional failure. longitudinal stresses are compressive at the 'core' of the wire and cause failure by longitudinal cracking or'splitting'. Longitudinal cracking seen at or 121 :\ V Spot llagn Dct WD Erp 200 pm kV 6.0 130r SE l¿1.6 6 f 0õ:lTact:*1 FIGURE 4.16 Deliberate fracture of 0.016 inch (0.406a mm) Wilcock'premium plus', batch TT-TAT. 122 I , {l '-v' t , rl:''- \ -' l.a>-' V Spot Magn Det WD ExP 100 pm 5 0 kV 6 0 200x SE 31.1 47 TT-TAT"deliberate#105G TC -G FIGURE 4.17 Detail of deliberate fracture of 0.016 inch (0.4064 mm) Wilcock'premium plus', batch TT-TAT. 123 '¿ '-î.-iú'/ - ,-7 - -t¡ T \< 50 pm Acc V Spot Magn Det WD ExP 15 fl lrv 6 fl 400x sE 316 49 TT-TAT:deliberate#1 FIGURE 4 - 1B Further detail of deliberate fracture of 0.016 inch (0.4064 mm) Wilcock'premium plus', batch TT-TAT. 'T' : Transverse narrowing in area of transition from cross-sectional to longitudinal failure 124 near the surface suggests that there is a compressive element to the sum of the near-surface residual stresses. Thinning of the wire across the transverse dimension suggests the influence of radial tensile stresses during fracture. These observations are consistent with those for wires drawn at successive reductions greater than one per cent, ie "the longitudinal stresses are tensile at the surface and compressive at the axis, the radial stresses are compressive at the axis, and the circumferential stresses follow the same pattern as the longitudinal stresses", (Dieter, 1961). In a single specimen bent at 50 degrees with tungsten carbide tipped light wire pliers, a split was seen. This longitudinal failure was approximately 200 micrometres long, ran parallel to the internal, concave surface, and was about one third the wire's thickness from the surface. There was no evidence of inclusions within or adjacent to the longitudinal failure that may be indicated as a point of failure initiation, however, the site of the failure seemed to be associated with the defect caused by the pliers at their point of application. The longitudinal failure is shown at FIGURES 4 - 19 and 4 - 20. Wiltiams, (1974) suggests that from observations of fractured specimens, the hardest, most work hardened part of the wire is at its sudace. His findings for micro-hardness confirm that this part and the innermost 'core' are the hardest areas. Both zones are about 40 micrometers wide. He suggests that greenstick fractures are initiated in the intermediate zone of uneven micro-hardness. The findings of the specimens in FIGURES 4 - 16 lo 4'20 would support this view. ln another specimen, a series of lines of etched pits was seen to extend perpe cu al e represent a series of lines of altered structure, possibly an increase in ductility, 125 -<--4'> ¿ V Spot Magn Det WD Exp 100 pm 20.0kV60 2ffix sE 116 7 TT-TAT:105:50deg:*4 F¡GURE 4 - 19 Longitudinal failure ('F') seen in specimen bent at 50 degrees with 'Dentronix 105' pliers Note the plier edge defect on the internal surface 126 -- --:--:---i Acc V Spot Magn Det WD Exp 50 pm 'tT-TATr105 æ0kV60 500x sE 117 9 50deg:14 FIGURE 4 - 20 Detail of longitudinal failure seen in specimen bent at 50 degrees w¡th 'Dentronix 105' pliers 127 and could be a pre-failure condition due to their ability to act as points of concentration for dislocations, (Dieter, 1961). The specimen is shown in FIGURE 4 - 21, and in greater detail in FIGURE 4 - 22. Where these lines of failure radiate inward from the external surface, any point stress applied to the wire surface during use, such as masticatory force at the point of contact between the wire and a bracket, is likely to increase the risk of failure. A failure that may result from such a change in structure has been reported recently in the literature, (Lee, 1992), and is illustrated at FIGURE 4 - 23. lt is important to note that the fracture described by Lee, (1992), stressed the role of excessive masticatory forces. The patient was able to apply a heavy shear force to the 0.014 inch, (0.365 mm) wire during mastication. Alternatively, such a fracture may be attributable to cyclic fatigue in the region of a surface irregularity, as described by Coquillet et al, (1979). All possibilities point to the significant contribution of clinical variables. Particularly of note is that this relatively fine diameter arch-wire had been in place for ten months. This would be considered an extended period for contemporary Begg treatment, (Jenner, 1993). During an extended period a finer than usual wire would be susceptible to a variety of oral fluids and excess mechanical cycling. These are the most likely causes of the fracture, irrespective of the microstructure. 128 FIGURE 4.21 Radiating lines ('R') of deformation that may represent a pre-failure condition. 129 F]GURE 4.22 Further detail of a radiating line of deformation that may represent a pre-failure condition' 'S' : External sudace of wire samPle 130 FIGURE 4 - 23 Exampleofatransversefailureofhightensilewire possibly resulting from transverse strain inherent in the wire, (Lee, 1992) (with kind permission from the Australian Orthodontic Journal) 131 TENSILE TESTING OF WILCOCKTM WIRES Wires were tensile tested as described in Chapter 3, and the tensile failure load result for each of six specimens collated to provide a mean failure load in kilograms. Given a knowledge of the diameter of each specimen, the mean load failure value for each series of specimens was inserted into the following equation to provide the tensile strength in Megapascals: TS= where: TS= tensile strength in MegaPascals, t-Þ: load in kilograms, 9= 9.8 metres per second squared, d= wire diameter in millimetres. The results obtained with the testing of wires supplied by A.J. Wilcock Scientific and Engineering Equipment are detailed in TABLE 4 - 3- For ease of comparison, results for tests performed by the manufacturer and the Australian Dental Standards Laboratory are shown in TABLE 4 - 4. Using those classifications detailed in the Australian Standard, AS 1964 - 1gZ7,the results shown in TABLE 4 - 4 would not classify all wires the same as has been done by the manufacturer. Similarly, results obtained in the current study indicate yet another classification for some of the wires' These discrepancies are indicated in TABLE 4 - 5. 132 Specimen Load failure in kilograms Diameter Batch T1 T2 T3 T4 T5 T6 Mean 0.307 mm HW-Rtr 19.3 1 9 1 19.8 19.1 19.2 19.2 19.3 0.358 mm DH-RIK 26.7 26.5 26.6 26.8 28.0 26.4 26.8 0.411 mm wr-Rco 39.1 41.4 39.2 39.2 38.9 39.3 39.5 * t t t 0.406 mm EH-RKL 41.4 41.7 41.6 0.513 mm RZ-ROI 65.8 62.3 62.0 61.8 62.3 61.5 62.6 TABLE 4.3 Results of each val¡d tensile test, (T1 to T6), and the resultant mean tensile load for failure for each batch of specimens. * Note: Sufficient wire available for 2 tests only Specimen Tensile Strength (103 MPa) Details Diameter Type / Batch AJW ADSL SLW 0.0121" Regular 2.397 2.324 2.550 (0.307 mm) HW-Rtr 0.0141" Reg. Plus 2.517 2.287 2.607 (0.358 mm) DH-RIK 0.0162" Spec. Plus 2.712 2.593 2.911 (0.411mm) WI.RCO 0.01 6" Prem. Plus 3.027 2.881 3.143 (0.406 mm) EH-RKL 0.0202" Spec. Plus 2.857 2.740 2.967 (0.513 mm) RZ-ROI TABLE 4 - 4 Tensile strength of Wilcock wires as measured by A.J. Wilcock Scientific and Engineering Equipment, (AJW), the Australian Dental Standards Laboratory, (ADSL), and exPeri mentation, (SLW). 133 Batch and A.JW ADSL SLW Diameter classification classilication classification IIW-RII Regular Regular Regular Plus 0.307 mm (Type 2) (Type 2) (Type 3) DH-RIK Regular Plus Regular Special 0.358 mm (Type 3) (Type 2) (Type 4) wI-RCO Special Plus Special Premium 0.411 mm (Type 5) (Type 4) (Type 6) EH-RKL Premium Plus Premium Premium Plus 0.406 mm (Type 6) RZ-ROI Special Plus Special Plus Premium 0.513 mm (Type 5) (Type 5) (Type 6) TABLE 4.5 Comparison of orthodontic wire classifications, as classified by .t ü A.J. Wilcock Scientific and Engineering Equipment, (AJW), Australian Dental Standards Laboratory, (ADSL), and during exPerimentation, (SLW). During calculations, the actual wire diameter was used to derive tensile strength results listed in TABLE 4 - 4. In order to test possible discrepancies, the nominal diameter, as distinct from the actual diameter, was substituted into the formula. This was considered a useful exerc¡se, s¡nce the value for the diameter was squared, and thus had the potential to alter the results significantly. The results are shown in TABLE 4 - 6 and compared with those results obtained for calculations using the actual diameter. 134 Actual (measured) diameter Nominal (labelled) diameter Batch (inches) Classification (inches) Classification * HW-RII 0.0121 Regular Plus 0.012 Special* DH-RIK 0.0141 Special 0.014 Special wr-Rco 0.0162 Premium 0.016 Premium EH.RKL 0.01 6 Premium Plus 0.016 Premium Plus * * RZ-ROI 0.0202 Premium 0.020 Premium Plus TABLE 4 - 6 Comparison of wire tensile classification as calculated with actual versus nominal diameters * Note: denotes a change in classification. 1 inch = 25.4 millimetres li ,l I I tt is noteworthy that the two changes in classification elevate both of the particular specimens into a higher category of tensile classification. This has the effect of furthering the discrepancies between tests performed during current experimentation and those results of A.J. Wilcock and the Australian Dental Standards Laboratory. tt is suggested that use of the nominal diameter in preference to the actual diameter is not val¡d, irrespective of the results obtained. Discrepancies in test results were noted by Ware & Masson (1976)' when testing Wilcock 'special plus'sg (Type 5) grade wire on behalf of the Australian Dental Standards Laboratory (ADSL). Original tensile tests by Masson, (¡b¡d) downgraded specimens of 'special plus' wire to 'special' (Type 4)' 33 A.J. Wilcock Scientific and Engineering Equipment, Whittlesea, Victoria, Australia ì 135 although subsequent tests in association with the ADSL showed consistent results commensurate with the manufacturers'classification. However, there was no specific indication as to whether tests had been carried out on the same specimens or if some 'age strengthening' of the wire had occurred. The text (¡b¡d) lacks clarity on this issue but probably indicates that separate specimen batches were involved. A similar series of tests on T.P. 'special Plus'g¿ (0.016 inch / 0.4064 mm - Orange spool) was conducted by Wiltiams, (1974) where discrepancies were found between different batches in both wire diameter and ultimate tensile strength. The range of diameter variation among eight separate spools of nominally 0.016 inch wire, was between 0.0158 inches and 0.0162 inches (ie a tolerance of 0.4064 mm * 0.005 mm). These diameters are within the specifications of the Australian Standard that stipulates a tolerance of t0.010 millimetres for wires with a nominal diameter of over 0.30 and up to 0.50 millimetres. The range of tensile values for these wires was from 2.65 - 2.7g x 103 MPa. These measurements indicate those specimens at the lower end of the tensile value range need re-classification as 'special' (2.551 lo 2.700 x 103 MPa) rather than 'special plus' (2.701 to 2.860 x 103 MPa). This re-classification is needed if wires are to conform to the Australian Standard, 4S1964 - 1977. Tensile test results are affected by various conditions in the wire, including microstructure and grain size. Additionally, certain factors associated with actual testing and the treatment of specimens are important in failure analysis if erroneous data are to be avoided. Variables to consider in testing include the strain rate applied to the material, since as strain rate increases so does lr 34 Manufactured by A.J. Wilcock, Australia, marketed by TP Orthodontics lnc., La Porte, lndiana, USA. 136 tensile strength, particularly at higher temperatures. For some materials this relationship may be critical even at room temperature. Generally however, as temperature increases, strength decreases and ductility increases, (Colangelo & Heiser, 1974). ln addition, other phenomena may occur, including the elimination of the yield point. Temperature effects can be compounded, depending on the physical metallurgy of the system and might include precipitation, strain ageing and recrystall¡sation. Other metallographic changes can occur with increased temperature, such as secondary hardening, where there is an increase in strength, or, a decrease in toughness, known as blue brittleness, (ibid). These effects, which are exposure time dependent, are not considered, since all tensile testing was conducted at room temperatures of between 20 and 25 degrees Celsius. An important point to consider is that much of the material cited in engineering texts in relation to mechanical testing, is for specimens much larger than the fine orthodontic wires currently under investigation. Those tests used in the mechanical engineering sphere are generally standardised in accordance with the requirements of the American Society for Testing Materials (ASTM). High tensite, cold drawn orthodontic wires have not been included in test parameters and as such it cannot be assumed that they will behave the same way as larger, less anisotropic materials. ln addition, the very small diameter of the orthodontic wires means that consideration of measurement error as a percentage of size might have a relatively greater effect on results. This could in part explain the consistently different results in tensile test data as obtained in this investigation, compared to the data from both the Australian Dental Standards Laboratory and A.J. Wilcock. 137 Tensile testing results in this investigation are consistently higher than those of A.J. Wilcock, which in turn are consistently higher than those of the Austratian Dental Standards Laboratory. TABLE 4 - 7 shows the percentage difference between the results for actual diameters. Part of the noted differences might be explained by the inherent inaccuracies of the 'bucket'test as used by A.J. Wilcock during initial wire classification. The momentum developed by weight addition in this test is inaccurate, and is equivalent to a constantly variable crosshead speed in a tensile testing machine. Since calculation of tensile strength is performed using deformation theory, ie relating stress to strain, validation of results can only be performed where the material is subjected to proportional loading, without any change in direction of the applied load, (Dieter, 1961). This problem is negated by the use of a constant crosshead speed, as specified in the Australian Standard, AS- 1964, (Standards Association of Australia, 1977). lt has been suggested that variations in crosshead speed may alter the measured tensile strength in some materials, (Cotangelo & Heiser, 1974)- Tensile testing performed in this investigation was carried out in accordance with the provisions specified in the Australian Standard, AS 1964. The specimens were tested on a machine operat¡ng at a crosshead speed of 0.5 t 0.1 millimetres per minute, and fitted with Hounsfield wire testing wrap around grips at least 50 millimetres apart. These variables are the same as those used by the Australian Dental Standards Laboratory. The variation in results between this investigation, (conducted in 1994), and the Australian Dental Standards Laboratory, (conducted in 1985), ranged from 7.g5to 12.72 percent, or, an average difference in the region of 10 percent. Given the effect this variation has on wire classification and the 138 Specimen (1Os Details Tensile strength MPa) Lol Diameter Type/Batch AJW ADSL Lto SLW f.o/o 0.01 21 " Regular 2.397 2.324 - 3.04 2.550 + 6.38 (0.307 mm) HW-RII 41 " Reg. Plus 0.01 2.517 2.287 - 9.14 2.607 + 3.58 (0.358 mm) DH-RIK Spec. Plus 0.0162" 2.712 2.593 - 4.39 2.911 + 7.34 (0.411mm) wr-Rco 0.016" Prem. Plus 3.027 2.881 - 4.82 3.1 43 + 3.83 (0.406 mm) EH-RKL 0.0202" Spec. Plus 2.857 2.740 - 4.10 2.967 + 3.85 (0.513 mm) RZ-ROI TABLE 4 - 7 Percentage differences in tensile strength as tested by the Austratian Dental Standards Laboratory, (ADSL), and exper¡mentation, (SLW) when compared to results of A.J. Wilcock Scientific and Engineering Equipment, (AJW). 139 resultant marketing implications, a clearer reason for the variation would be useful. The tensile classification of wires has a profound clinical implication, in that without a true knowledge of the actual wire characteristics, the forces being applied to teeth in patients' mouths are potentially unknown. Variation and inconsistencies in tensile tests are common, (Williams, 1974), such that ultimate tensile strength is measurable with great accuracy, whereas 0.1 per cent proof stress and proportional limit are less well defined. Without the use of an extensometer, tensile tests are only useful for general comparison of wires, (Williams, 1974; Allen 1994). This also leads to inaccuracies in determining the modulus of elasticity. The issue of age changes may be involved in differences found between tests carried out on the same batches of wire nine years apart. Reliable information on ageing mechanisms has not been achieved for heavily drawn steel wire with atensile strength in excess of 2x 103 MPa, (Davies, 1987). The observed rise in the ultimate tensile strength may be a factor of strain ageing that is said to occur following an interruption in the deformation process, such as experienced by storing the wire for extended periods prior to use, even at room temperature, (Honeycombe, 1981). Strain ageing may affect other physical properties such as tensile strength, yield stress, fracture toughness, and ductility, (Davies, f 987). Strain ageing would seem to result from the migration of carbon atoms within the grain structure. Once the carbon atmospheres have pinned the dislocations within the wire structure, they remain locked in 'Cottrell atmospheres'. This seems to be particularly the case within ferritic phases, (¡b¡d). Solutes that have a high diffusivity, such as nitrogen, oxygen or boron, may also pin 140 dislocations, and contribute to strain ageing. Various alloying elements may alter the solubility of carbon and nitrogen, and thus alter the pinning capacity and strain ageing ability of these solutes. This can be done commercially either by reducing the concentration of interstitial atoms such as carbon or nitrogen, or, by the addition of alloying elements that encourage the formation of carbides and nitrides and 'tie up'the excess interstitial sotutes. Nickel will increase the strain ageing in steels, whereas manganese and phosphorus may slow down low temperature strain ageing, (¡b¡d). Sudden generation and movement of newly formed dislocations and the formation of precipitates on the dislocations, as arises when clinically manipulating 'old'wire, lead to a return in the yield point. This phenomenon occurs at the cost of a higher dislocation density, and thus greater risk of fracture, because of an adverse effect on the wire'S ductility, (Honeycombe, 1981; Davies, 1987). The control of solute content during manufacture, and thus carbon and nitrogen atom diffusivity during storage and clinical use, would seem important in this respect. Storage of Wilcock wires for extended periods would therefore seem contra-indicated. Defining this 'extended period' cannot be quantified from the evidence of this investigation. Results hitherto discussed suggest that sufficient doubt exists to quote tensile data accurately on commercially available high tensile wire labels. This practice has occurred in the past because "ultimate tensile strength is... well defined in atensiletest... (and is) the only parameterwhich may possibly be related to elast¡c properties (in the application of Industrial Standards)", (Twelftree, 1974'). The use of the wire's proof stress might be a more reproducible quantity for the assessment of elastic properties, (¡b¡d). Williams, (1974), suggests that force at proof, which is the absolute force 141 that a wire may withstand without permanent distortion, has greater clinical significance than proof stress, as suggested by Twelftree' (1974). The highlighted inconsistencies in determining tensile data for high tensile wires discussed thus far may shed light on the inconsistencies for published values of Young's Modulus of elasticity as discussed by Wiltiams, (19741, Goldberg et at, (1983b), Yoshikawa et al, (1981) and Allen, (1994). This may reflect differences in experimental technique, particularly where there is inconsistency for values being derived from either tensile or flexural data. However, it has been suggested that for stainless steel w¡res, the tensile modulus of elasticity is very similar to the flexural modulus of elasticity, since the tensile and compressive properties are very similar, (Yoshikawa et al, 1981). Further, it has been suggested by Masson, (1969)' that differences in values of elastic modulus for stainless steel wires are dependent on the wire being assessed. Differences in the amount of cold drawing, as is seen in the production of varying tensile grades of wire, will show a commensurate difference in the measured values for Young's modulus of elasticity. Thus for a higher tensile value, a wire can be assumed to have a higher value for its modulus of elasticity. Where there is doubt as to a wire's true tensile strength, so there must also be doubt as to its true value for modulus of elasticity. Williams, (1974), suggests on the basis of his study, that the clinical acceptability of orthodontic wires cannot be demonstrated by any tensile method' He favours the mandrel testing method of Stephens & Waters, (1971), as being clinically acceptable, and, the use of Young's modulus of elasticity, as it is a measure of flexural rigidity and thus can be related to the intended clinical use. 142 ARCH-FORM ANCHORAGE BEND ANALYSIS ln the clinical situation, when an arch-wire is engaged into brackets on irregularly aligned teeth, the wire is deflected to differing radii of curvature at different points around its arc. As such, the initial activated arch-wire forces, incident upon the teeth in an individual clinical situation defy analysis, (Burstone & Koenig, 1974). A weakness of any apparatus designed to simulate and measure arch-wire forces cannot take any account of irregularities in tooth position in all three planes of space, nor can it account for the dynamic effect on force levels that must occur as teeth move. Given this proviso, it was decided to observe the force levels at different points around a typical Begg stage 1 arch-wire. The apparatus chosen in some part surpasses the systems' analytical complexity, in that stored energy, or strain energy available at a particular point or points on the arch-wire, can be measured at the point of application. The major limitation of the model arch-wires was their departure in form compared to the clinical archwires they were meant to represent. This departure was particularly so in relation to the relative lack of posterior expansion and distal toe angle. However, it is considered reasonable that within the limits of those constants discussed in 'methods', observations made in respect to the experimental apparatus could be related to clinical behaviour of the wire. Alterations in measurable force levels will occur depending on the type of beam support, (Nikolai, et al, 1988). Adams et al, (1987)' showed that this difference applies when different types of bracket are used, and, with the method of ligating the wire to the bracket. ln addition, the force levels will 143 vary depending on the presence and type of adjacent brackets. lt was decided to observe the experimental forces using a point support as this was considered to approach the clinical behaviour of ribbon arch brackets, (Swain & Ackerman, 1969), as used in the Begg technique. To consider an orthodontic arch-wire as a basic point supported beam is simplistic, (Drenker, 1988). However, it serves as a useful basis to understanding the effects cenain parameters will have on the forces present. For a cylindrical beam in two dimensions held between two supporting points, the amount of deflection with the application of a load is based on the formula: 4.P.L3 D = g.ß.d\E Rearranging this formula to obtain an expression for the effect of changes in load, is as follows: where: l-Þ: load applied between the suPPorts, D= deflection of the beam supported at both ends, L_ L_ Young's modulus, d: wire diameter, t_ L- beam length between the suPPotts, (Greener et al, 1972). These formulae suggest that within the working range of an orthodontic wire, the variables of wire diameter and length of span will have the greatest effect on toad - deflection characteristics. This is in keeping with engineering principles. The essence of this observation will also be true for a three dimensional arch-wire, (Rohan, 1982). 144 It might be assumed that in the analysis of arch-wire mechanics, an intrusive force applied by a Stage 1 Begg arch-wire will have a similar magnitude of force on all those anterior teeth engaged in the arch-wire. However, the work of White et al, (1979), and later of Murphy et al, (1982), clearly showed that with a bioprogressive contraction utility arch, the initial force is principally borne by laterat incisors, that is, by the most laterally engaged teeth. ln Begg stage 1 mechanics, (Begg & Kesling, 1977), this would generally apply to the canines. These findings are extant only in an ideal situation where the teeth are aligned; irregularities of the teeth in all three planes of space, as occurs in most individuals, would lead to a vast envelope of possibility. Each specimen was vertically loaded and force levels recorded after 60 seconds. This allowed for the 'initial transient', as discussed by Hazel et al, (1984), where it was shown that for premium plus wires, stress relaxation accounted for up to a 10 per cent reduction in force levels during the first 60 seconds. Compared to other available Wilcock wires, the 'premium plus' grade maintains the greatest residual force levels over time, and the lowest level of stress relaxation, (¡b¡d). lt is likely that the new 'supreme'35 grade wire exceeds the stress relaxation advantages of 'premium plus' wire. Noting that stress relaxation is a dynamic phenomenon, (Morris et al, 1981), and obeys a logarithmic decay rate, (Hazel et al, 1984), a transient of 60 seconds was considered clinically valid for measuring 'initial forces'. The measurements for each degree of anchorage bend, as a function of the load position are listed in FIGURES 4 - 24 lo 4 - 27. Additionally, the position of the 'knife-edge' support in each arch-wire type is represented in linear, and logarithmic graphical form. 35 A.J. Wilcock Scientific and Engineering Equipment, Whittlesea, Victoria, Australia 145 levelat different locations, (grams) - 15" bends Arch-wire Force type Cenlre lncisal 2's 3's 4's Plain 16.25 17.78 20.58 41.22 1 53.1 4 Circles 14.30 16.28 18.35 36.30 140.02 Boot hooks 15.01 16.74 19.44 39.92 144.24 Loop hooks 14.51 16.15 18.76 34.78 130.09 Molar stops 15.15 16.65 19.23 35.46 112.54 160 140 120 -*-Plain 100 --1O-circles Force hooks (grams) BO {-Boot 60 ' ' ri , "'Loop hooks 40 Molar stops 20 0 Centre lncisal 2's 3's 4's Knife-edge position on arch-wire 1 000 Pla in 100 Force #Circles log 10 --{-Boot hooks (grams) '10 :'l: Loop hooks Molar stops 1 Centre lncisal 2's 3's 4's Knife-edge position on arch-wire FIGURE 4.24 Analysis of test arch-wires with 15 degree anchorage bends. 146 different locations, (grams) - 30" bends Arch-wire Force levelat type Centre lncisal 2's 3's 4's Plain 29.50 35.57 42.37 95.1 2 341.61 Circles 28.64 34.34 41.18 92.72 344.31 Boot hooks 30.47 35.44 44.27 93.34 334.56 Loop hooks 29.03 33.97 41.68 82.28 334.21 Molar stops 29.29 33.95 41.76 80.38 291.74 3s0 300 Plain 250 *Circles Force 200 $Boot hooks (grams) t S0 -;r:. Loop hooks 100 Molar stoPs 50 0 Centre lncisal 2's 3's 4's Knife-edge position on arch-wire 1 000 --C-Plain 100 Force #Circles log 10 {-Boot hooks (grams) 10 ;ri-i " Loop hooks Molar stops 1 Centre lncísal 2's 3's 4's Knife-edge position on arch-wire FIGURE 4 - 25 Analysis of test arch-wires with 30 degree anchorage bends. 147 Force levelat different locations, (grams) - 45o bends Arch-wire type Centre lncisal 2's 3's 4's Plain 47.18 54.77 60.40 147.47 404.16 Circles 42.45 49.29 54.98 134.88 398.35 Boot hooks 45.10 53.23 59.26 151.44 419.47 Loop hooks 41.60 48.28 55.27 134.30 390.90 Molar stops 39.12 43.35 51.28 134.72 306.28 450 400 350 Pla in 300 *Circles Force 250 hooks (grams) 200 *Boot 150 :)<,. Loop hooks 100 Molar stops 50 0 Centre lncisal 2's 3's 4's Knife-edge position on arch-wire 1 000 *Plain 100 Force *Circles log 10 hooks (grams) -¡!-Boot 10 ;:,,. Loop hooks Molar stops 'l Centre lncisal 2's 3's 4's Knife-edge position on arch-wire FIGURE 4 - 26 Analysis of test arch-w¡res w¡th 45 degree anchorage bends. 148 different locations, (grams) - 60'bends Arch-wire Force levelat type Centre lncisal 2's 3's 4's Plain 66.10 75.07 92.75 155.13 479.02 Circles 60.21 71.36 88.05 139.58 454.81 Boot hooks 61.81 72.66 88.49 141 .08 442.14 Loop hooks 62.13 74.22 88.09 144.65 438.26 Molar stops 58.13 69.05 83.19 145.01 423.67 s00 400 Pla in -#Circles Force 300 Boot hooks (grams) f 200 ';. i' 'Loop hooks 100 Molar stops 0 Centre lncisal 2's 3's 4's Knife-edge position on arch-wire 1 000 Pla in 100 Force #Circles log 10 hooks (grams) -¡!-Boot 10 '., ,'' Loop hooks Molar stops 1 Centre lncisal 2's 3's 4's Knife-edge position on arch-wire FIGURE 4 - 27 Analysis of test arch-w¡re with 60 degree anchorage bends. 149 Hooke's law defines the relationship of wire deflection to applied force, such that w¡thin the elastic limit of a material, stress is proportional to strain. Therefore, throughout the range of arch-wire activation, the quotient of applied force divided by deflection is a constant, termed the load - deflection rating, (phittips, 1923). This rating will vary with different arch-wire construction, such as with the addition of inter-maxillary hooks, or different anchorage bends. A wire with a high load - deflection rating is considered preferable, since it can act continuously between adjustments. Similarly, if the wire also has a good working range, then the number of adjustments required are reduced. These factors allow efficiency in treatment, (Rohan' 1982). Munday, (1969), summarises those factors that govern anchorage bend severity, and thus in part, arch-wire function, in Begg mechanics, namely: a the d¡ameter of the wire, a the distance of the anchorage bend from the canines, (3's), a the cant of the buccal tubes, o the relative tilt of the molars, o the curve of the occlusal plane, or curve of Spee, a the amount of toe-in or toe-out, a the length and diameter of the buccal tubes, a the axial inclination of the anterior teeth. tn contrast, Sampson, (1992), considers treatment variables that determine anchorage bend severity, such as the growth status and dento-facial type of the individual, the anchorage requirements, intrusion requirements, arch-wire dimensions, arch-wire grade, anchorage bend position, the presence of hoOks, loopS, stops, etc, and, the stage of treatment. DeCisions on 150 anchorage bends are based on these factors and an assessment of treatment objectives, that is, the harmonious positioning of the teeth by orthodontic therapy within the stomatognathic system. Forces required to intrude anterior teeth have been variously described by a number of authors. The force values quoted range from 15 to 25 grams per tooth, (Proffit, 1993), 40 grams for a six tooth incisal segment, (Thornton & Nikotai, 1981), and 200 grams for an incisal segment, (Burstone, 1977; Ricketts et al, 1979). There have been similar variations in the recommendations for anchorage bend severity. Begg & Kesling, (1977) suggest that the anchorage bend should allow the anterior part of the Stage 1 arch-wire to lie passively in the muco-gingival sulcus rather than specifying a bend of a known degree, but suggest a bend of 30 degrees as a guide. Thompson, (1972) and Baldridge, (1973), suggests the passive wire should lie deep in the muco- buccal fold. Alternatively, Sims, (19771 suggests that adequate intrusion from Stage 1 arch-wires can only be achieved with anchorage bends greater than 30 degrees and suggests 45 degrees. An examination of the information presented in FIGURES 4 - 24 lo 4 - 27 shows a number of interesting points. The first, and perhaps surprising is the similarity of results between arch-wire types. Despite the differing forms of the arch-wires, the anterior intrusive forces were remarkably similar for a given degree of anchorage bend. This is despite some inter-maxillary hooks being bent in different planes, as is the case when comparing circles, bent in the horizontal plane, against boot and loop hooks, bent in the vertical plane. 151 Where no inter-maxillary hook had been bent into the arch-form, as in the 'plain'arch-wires, the only significant difference was noted in the 15 degree sample. The 'plain' arch-wire exerted an intrusive force approximately 10 per cent greater than those arch-wires with some form of inter-maxillary hook. Conversety, the 15 degree arch-wire with molar stops exerted an anterior intrusive force that was approximately 10 per cent less than those arch-wires with inter-maxillary hooks. These observations are in keeping with load - deflection relationships that arise as a result of the existence and position of the increased wire length incorporated in either an inter-maxillary hook or molar stop. The finding of a decreased anterior intrusion capability for the 'molar stopped' arch-wire was noted for all degrees of anchorage bend, particularly in the 4's, (first premolar) location. Given that the distance from first premolar to the anchorage bend was the smallest of the distances investigated, the reduced anterior force levels for 'molar stopped' arch-wires are in accordance with an increased length of wire, (Newman, 1963), since the load - deflection ratio of the wire is inversely proportional to the cube of the length of wire between supports, (Greener et al, 1972\. When studying double-back 0.016 inch (0.4064 mm) arch-wires in oval tubes, Wiltmot, (1983) noted that adding uptake loops to the arch-wires had a marked reduction in the intrusive capacity of arch-wires. The intrusive capacity of the looped, double-back arch-wire was less than for a plain double-back arch-wire with an increase in the degree of anchorage bend. This is probably indicative of a progressive 'opening' of the uptake loop during increased elastic deformation as a function of an increase in the degree of anchorage bend. 152 From observation of the results tabulated in FIGURES 4'24 to 4 - 27,4t appears that a significant difference exists in intrusive force capability when the first premotars are engaged in an archwire. For a given degree of anchorage bend, the force tevels at the premolar region are consistently 10 times greater than in the anterior region. Perusal of the logarithmic graphs illustrates this point. This finding is not as clear for the arch-wires with 60 degree anchorage bends. tt is likely that the significant rolling of the wire in the molar buccal tubes for these wires had a marked distortion on some results. Not withstanding this, the fact that wire rolling was observed in the molar buccal tubes, is suggestive of forces in excess of those considered as biologically acceptable, (Brezniak & Wasserstein, 1993), and detrimental to the efficiency of the Begg appliance, (Sims, 1977). Wiltmot, (1983) surmised that Class 2 elastics would reduce the intrusive capacity of a maxillary arch-wire. Recent attempts to quantify the intrusive force reduction effect of Class 2 elastics have been published by Xu et al, (1gg2). Using a stepwise regression analysis of results, the authors found less reduct¡on in the intrusive capacity of the arch-wire than had been previously believed, (¡b¡d). The impoftance of understanding the dynamic balance of moments and couples in archwires, and the effect of inter-maxillary elastics during treatment, will provide an insight ¡nto the requirements of anchorage bends in the Begg appliance, (Sims, 1977; Hocevar, 1981). To gain a better understanding of how each of the arch-wire forms behaved as a function of the degree of anchorage bend, the results from FIGURES 4 - 24 to 4 - 27 have been re-arranged. By so doing, the results could be compared with the results of other cited authors. The results, in tabular and graphical form, are shown in FIGURES 4 - 28 to 4 - 32- 153 (grams) - plain arch-wires' Anchorage Force level at d¡fferent locations, Bend Centre lncisal 2',s 3's 4's 15 deg '16.25 '17.79 20.58 41.22 1 53.1 4 30 deg 29.50 35.57 42.37 95.1 2 341.61 45 deg 47.18 54.77 60.40 147.47 404.16 60 deg 66.1 0 75.07 92.75 1 55.1 3 479.02 Plain arch-wires s00 400 300 Force -Çs¡t¡s***-*-:.--*2',s (grams) 2g9 -lncisal*-'- '* '-3's 100 0 15deg 30 deg 45 deg 60 deg -{'g Degree of anchorage bend FIGURE 4 - 28 Analysis of 'plain' arch-wires with vary¡ng degree anchorage bends. 154 't. locations, (grams) - inter-maxillary circle arch-wires. Anchorage Force level at different Bend Centre lncisal 2's 3's 4',s 15 deg 14.30 16.28 18.35 36.30 140.02 30 deg 28.64 34.34 41 .18 92.72 344.31 45 deg 42.45 49.29 54.98 134.88 398.35 60 deg 60.21 71.36 88.05 139.58 454.81 lnter-maxillary circle arch-wires s00 400 Force 300 (grams) 200 100 0 15deg 30 deg 45 deg 60 deg Degree of anchorage bend F¡GURE 4 - 29 Analysis of inter-maxillary circle' arch-w¡res with varying degree anchorage bends. 155 locations, (grams) - boot hook arch-wires. Anchorage Force level at different Bend Centre lncisal 2's 3's 4',s 15 deg 15.01 16.74 19.44 39.92 144.24 30 deg 30.47 35.44 44.27 93.34 334.56 45 deg 45.1 0 53.23 59.26 151.44 419.47 60 deg 61.81 72.66 88.49 141 .08 442.14 Boot hook arch-wires s00 400 J ll Force 300 -centre*----.---'---*--= 2 's (grams) 2gg -lncisal-.'-*.-*-3 ' s 100 0 15deg 30 deg 45 deg 60 deg -Q'3 Degree of anchorage bend FIGURE 4 - 30 Analysis of 'boot hook' arch-wires with varying degree anchorage bends. I t 156 locations, (grams) - loop hook arch-wires. Anchorage Force level at different Bend Centre lncisal 2's 3's 4's 15 deg 14.51 16.15 18.76 34.78 130.09 30 deg 29.03 33.97 41.68 82.28 334.21 45 deg 41.60 48.28 55.27 134.30 390.90 60 deg 62.1 3 74.22 88.09 144.65 438.26 Loop hook arch-wires s00 400 300 Force -Centre---l'g (grams) 200 -lncisal-'*-*-'* 3's 100 0 15deg 30 deg 45 deg 60 deg -{'5 Degree of anchorage bend F]GURE 4. 31 Analysis of 'loop hook' arch-w¡res with varying degree anchorage bends. 157 locations, (grams) - molar stopped arch-wires' Anchorage Force level at different Brnd Centre lncisal 2's 3's 4's 15 deg 15.15 16.65 19.23 35.46 112.54 30 deg 29.29 33.95 41.76 80.38 291.74 45 deg 39.12 43.35 51.28 134.72 306.28 60 deg 58.1 3 69.05 83.1 I 145.01 423.67 Molar stopped arch-wires 500 400 Force 300 (grams) 296 100 0 15deg 30 deg 45 deg 60 deg Degree of anchorage bend FIGURE 4 - 32 Analysis of 'molar stopped' arch-wires with vary¡ng degree anchorage bends. 158 The first point to note is the subtle force level differences between arch-wires at given support positions. This suggests subtle differences between archwires, particularly since each was individually fashioned. Support position had been etiminated as a variable by fixing the supports at each position for each series of tests. Examination of the data reveals a slight increase in force levels for the 60 degree bend in the plain arch-wire supported at the first premolar, (4's), region. However, as this difference is of the order of 5 per cent, it is likely to be experimental error, probably due to rolling of the wire in the molar buccal tubes. More important is the effectively linear behaviour of all wires when activated by supports in the incisor region. This is in keeping with the general findings of Borghesi, 1 973; Sims, (1977); Thornton & Nikolai, (1 981); Rohan, (1982); and Wiltmot, (1983). When positioning the 'knife-edge' supports in the canine, (3's), region, all arch-forms showed an initial linear behaviour with increases in severity of anchorage bend. However, a reduction in expected force behaviour, as illustrated by a loss of linearity in the graphs, was noted for the 60 degree bends. This is probably also related to the arch-wire rolling in the molar buccal tubes during experimentation. The behaviour of all arch-forms when loaded at the first premolar, (4's), region shows a consistent pattern for the 15 and 30 degree anchorage bends. When a load is applied with 45 degree anchorage bends, a drop in expected force level occurs, suggestive of arch-wire rolling. This occurs at an earlier stage when compared to wires loaded at the canine, (3's), position. Other than the similarity of results for incisal intrusive force levels, results from this investigation are not in agreement with the authors cited above. This disagreement is in relation to the forces exerted by 60 degree bends. Of 159 these authors, Rohan, (1982), is the only individual to agree with a reduction in expected linearity for arch-wires not clamped at the distal ends. The exptanation for this might relate to arch-wire rolling in the molar buccal tube, as atready described. lf anchorage bend toe angle had been finely adjusted for every arch-wire activation in each series of experiments, rolling may have been avoided and linearity of results might have been achievable. However, having fabricated the arch-wires it was decided to maintain their shape, since, all experimental behaviour was to be observed in the 'as-bent' form. ln this way no further variables were introduced into the study. Borghesi, (1973), has suggested that purely linear results for variation of anterior intrusive force as a function of anchorage bend severity can be achieved by clamping the ends of the arch-wire rather than using a molar tube. This has two limitations. Firstly, there is no freedom to the distal ends of the arch-wire, although the full potential of the anchorage bend can be measured. Secondly, those torsional changes that occur, particutarly with an increase in anchorage bend severity, cannot be observed. While arch-wire rolling is undesirable in clinical practice, there is an obvious need to understand its implications if it occurs. Where, however, the full potent¡al of the anchorage bends is not realised, those causative factors might be identified and understood. Finally, an assessment of forces that exist during each activation must include those forces and moments acting on the molar buccal tubes. Cadman, (1975), suggests that anchorage bends exert a distal tipping force on molars. This is often minimised by the multi-rooted anatomy of molar teeth and the forces of occlusion resisting such forces. Resistance may also be achieved by the use of 'Class 2' elastics, (Cadman, 1975; Hocevar' 1981; Thornton & Nikola¡, 1981; Xu et al, 1992), which act to counter the 160 moment of the lower molars as produced by the anchorage bend, (Mulligan' 1979, 1980; Burstone & Koenig, 1988; Ronay, 1989; Bequain' 1gg2). The results of this investigation would suggest that as the measured intrusive force at the anterior part of the arch-wire changes, so there is a commensurate change in the moment act¡ng on the molar via the molar buccal tube. This moment will undoubtedly have a component of transverse force depending on the degree of rolling that might occur with arch-wire activation. Thus, three dimensional control of molar positioning forces and moments is intimately related to arch-wire forces as derived from anchorage bends, (Cadman, 1975i Sims, 1975; Begg & Kesling, 19771. No attempts were made in this investigation to determine a mathematical relationship for arch-form anterior intrusive forces as a function of the d¡stance of the knife-edge supports from the molar tube. While this is straight-forward for a simple beam in two dimensions, the complexity of vectorial calculations involving beam theory in three dimensions was considered beyond the scope of this project. Attempts have been made to derive a formula that describes an arch-form, (Borghesi, 1973; Rohan, 1982), however, the analyses presented are an approximation at best. Three dimensional analysis of orthodontic mechanisms has been attempted, (Koenig et at, 1984), using mathematical transfer matrices. This system has a major flaw in attempting to understand arch-wire forces, since no consideration can be given to the constantly dynamic events of orthodontic tooth movement. Not only do the teeth themselves move, but as changes to the arch-wire configuration are needed progressively through treatment to cater for the tooth movement, so the arch-wire forces must change. A system of strain gauges between bracket and tooth might provide a method of monitoring such changes in terms of applied forces. 161 CHAPTER 5 CONCLUSIONS The 'premium plus' grade of 0.016 inch, (0.4064 mm) Wilcock wires are manufactured form AlSl Type 302 bulk wires. The resulting product is heavity drawn, resulting in a substantially cold worked structure. Much of the predictable structure of the wire relates to three distinct longitudinal layers; a thin outer'skin' zone that is about 40 pm wide, an inner core that is about 25 to 35 pm wide, and the remainder being an intermediate zone of uneven micro-hardness, and thus probably of varying mechanical properties. The extent and characteristics of these zones are influenced by the drawing process and heat treatment used during the manufacture of the wire. Etching regimes for microscopy may have variable results, particularly where there may be unknown variations in the preparation of specimens. Alternatively, although less likely, variations in the structure of the wire may account for some differences in the etched result as viewed with scanning electron microscopy. The use of potentiostat etching, after Williams & Fraunhofer, (1971), is likely to allow greater control over method variation. During the clinical manipulation of the wire, excess plier holding pressure needs to be avoided, as scanning electron microscopic evidence supports the hypothesis that this will damage the wire, and may increase the risk of 162 fracture. This risk is substantially increased with the use of tungsten carbide tipped light wire pliers that have a keen edge. When the wire was bent with tungsten carbide tipped pliers, the plier edge defect was seen to affect the structure of the wire 2 to 3 times as severely as the stainless steel light wire, ('139'), and Tweed torquing, ('442'), pliers. The defect continued through the harder outer zone, or skin of the wire and into the intermediate zone, an area where the initiation of 'greenstick' fractures is considered to occur. Keen edged, tungsten carbide tipped pliers cannot be recommended for routine wire bending as a result of these findings. The process of wire fracture might be as follows. A higher than normal inclusion density or dislocation density, may cause an increased potential for failure in a particular segment of wire. During clinical manipulation, this segment may undergo a sudden increase in stress and strain in the intermediate zone. This is particularly likely if an internal surface 'plier edge defect' occurs because of an excess 'plier holding pressure', or, if tungsten carbide tipped pliers are used. The result is an intermediate zone split. lf manipulating the wire also causes excessive strain in the 'skin' zone on the outer sudace, strain failure may occur. This will happen if there is a point at which the stress exceeds the tensile strength of the wire. Consequently the wire fails with a typical 'greenstick' fracture. Using a modified pair of tungsten carbide tipped pliers, after Mollenhauer, (lg9O), wire bending could be performed without causing any damage to the wire surface, as viewed with scanning electron microscopy. The major short- coming of these 'dolphin' pliers was the inability to bend conventional inter- maxillary hooks, (other than 'loop hooks'), due to the plier beak width. 163 TENSILE TESTING OF WILCOCKTM WIRES The composition of Wilcock wire is essentially austenite, however sufficient evidence is available to support the existence of õ-ferrite in small quantities. Very small quantities of metastable deformation-induced martensite may also exist. The presence of metastable marlensite formed in this way may change with different wire grades from variations in cold working, and the amount of heat treatment during manufacture. The likely structure or the wire has been described in the literature, and partly observed in this investigation. Age changes in the microstructure are considered to occur and to have an effect on measurable physical properties. Microstructural changes might be inferred from comparing tensile testing results for the same batches of wire that had been originally tested 10 years before this current investigation. Results were compared for the tensile testing of a series of different grade and diameter wires on three separate occasions. The first tests were following production, and tested by the manufacturer, A.J. Wilcock, in 1983 - 84. The second testing was by the Australian Dental Standards Laboratory, (ADSL), following submission of the wires in 1985. The final testing was in 1994 as part of this project. Comparison of the results of these three series of tests suggests that some change in the wire occurs over time, and, that there is a consistent variation in testing technique, since inter-operator test results were consistently differe nt. The results of the ADSL tests were consistently lower than those of the manufacturer, whilst the results of tests performed as part of this study were consistently higher than those of the manufacturer. lt is considered that the test results of the ADSL and this study are comparable, as they were both 164 performed to the specifications of the Australian Standard. The tests performed by the manufacturer were done as 'dead weight' tests, that, while reasonabty accurate, did not use a constant 'cross-head speed' and thus may not be sufficiently accurate to classify the wires correctly according to the Australian Standard. lt has been suggested that the manufacturer is soon to adopt the practice of wire testing with a tensite testing machine, (Wilcock, 1 ee4). The most plausible reason for the consistently increased results for ultimate tensile tests in this investigation is the effect of strain ageing. Despite an increase in the ultimate tensile strength, other properties such as formability will have suffered detrimental changes. The storage of Wilcock wires for extended periods cannot be recommended from the results of this investigation. Given the uncertainties of tensile testing results, validating values for Young' modulus, (E) is difficult. To this end, the use of a wrapping test, after Stephens & Waters, (1971), may be preferable. Force at proof, or, the maximum force applicable prior to plastic deformation, is a more clinically meaningful measurement for labelling purposes, (Williams, 1974). 165 ARCH.FORM ANCHORAGE BEND ANALYSIS The efficacy of Wilcock wires in their application to the Begg philosophy of orthodontics is due in the main to their high stiffness and excellent springback properties. The results from arch-form testing suggest that force levels exhibited by engagement of anterior brackets into an arch-wire are complicated, but predictabte. However, as arch-wire engagement continues distally, force tevets become tess predictable. The same problem arises more anteriorly as the anchorage bend degree is increased in severity. There are clinical implications for a Begg stage 1 arch-wire. These include a need for caution when considering engaging first premolar teeth into an arch- wire. Where true intrusion of the anterior teeth is fundamental to treatment Success, first premolars should be omitted from the active appliance, or anchorage curves used, as described by Lew, (1990). lf considerable incisal intrusion is required, consideration should be given to the omission of canines from the initial active arch-wire. This may occur in severe deep bite cases requiring anterior intrusion. Similarly, if moderate to severe anchorage bends are required for treatment needs, extreme care should be taken to adjust the 'toe' angle of the anchorage bend so as to prevent rolling of the wire in the molar buccal tubes. Visually checking the behaviour of anchorage beds in the activated state should be routine prior to ligating the arch-wire into the brackets. 166 tt can be concluded that anchorage bends in the Begg technique can provide light continuous anterior intrusive forces that are a function of the degree of anchorage bend, and of the teeth ligated to the arch-wire. As hypothesised by Thornton & Nikolai, (1981), with no torque potential in the appliance, the teeth may seek out the intrusive path of least resistance. Similarly, with long moment arms, even a light force against the anterior segment will effectively cause a large tipping moment to the banded molars. With larger anchorage bends, the potential for effective anterior intrusive force is reduced and transposed as transverse forces acting on the molar tubes and most posteriorly bracketed teeth. This is commonly referred to as 'arch-wire rolling'. Experimental data produced in this study may be at variance with other reported studies, and with future studies. The effect of different arch-forms and dimensions will be instrumental in this difference. 167 CHAPTER 6 FUTURE RESEARCH Williams, (19741 suggests that no investigations have looked at the possibility of the surface of a wire being more work hardened than the interior as a result of the drawing process. Variations in micro-hardness across the width of the wire would reveal possible differences. This would require a micro-hardness testing machine and may shed light in reasons for the observed clinical behaviour of Wilcock wires. Allen, (1994) suggested the use of an ultra-micro-indentation system. Any variation in grain structure across the range of Wilcock wires might be surmised from microhardness values. Using a micro Vickers test, Williams' (1974) found that for 0.016 inch (0.4064 mm) T.P.36 'special plus', microhardness showed the least variation in the peripheral 40 micrometre 'skin'zone and in the 40 micrometre core zone. Greatest variation was seen in the intermediate zone. Atternatively, the phase characteristics of the higher tensile wires might be established by ferromagnetic measurement and electrochemistry. The testing of wires aged in oral conditions may provide interesting information as to their obserued behaviour 36 T.P. orthodonl¡cs, lnc., La Pofle, lndiana, USA. 168 TENSILE TESTING OF WILCOCKTM WIRES Given that a number of authors have agreed on the lack of clinical applicability of tensite testing, performing an alternative set of tests on Wilcock wires, including the wrapping test might provide some useful information. Gaining access to other samples of aged wire and submitting them to tensile testing with an extensometer may give a truer indication as to some of the mechanical properlies that were unable to be measured in this study. This would include a consideration of the effects of strain ageing of Wilcock wires. Williams, (1974) claims that electro-polishing the high tensile Wilcock wires, as is the custom of some operators, should be avoided, as it removes the outer hard cold worked layer. A comparison of the microstructual and failure characteristics of 'as received' and electro-polished wires would settle this debate. 169 ARCHFORM ANCHORAGE BEND ANALYSIS Further studies are possible with the jig used in this study. The various combinations of arch-wire form and function are endless. The parameters of any future study are boundless. Applicability to a particular clinical problem in hand may provide useful information as to its solution. An alternative method of studying arch-wire forces incident on teeth might be to incorporate micro strain gauges beneath orthodontic brackets. Should a sufficiently accurate way be found to achieve this, such experimentation might be applied to an actual clinical situation. This would allow observation of the dynamic nature of orthodontic forces as the teeth move. Any increase however, in the distance between the arch-wire and the tooth's buccal surface will incorporate an altered balance between forces and moments, as described by Kusy & Tulloch, (1986), and extrapolation to the true clinical situation would be guarded. Arch-wire forces can be measured from a jig such as used in this study, however this sheds little light on how the teeth react within the alveolus on an individual basis for a given force system. An analysis of arch-wire function simitar to the clinical situation would be augmented by the use of photoelastic models. 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Res., 60(2): 139 - 145. 185 APPENDIX 1 PHOTOELASTIC MODELLING Photoetasticity can be used to study the stresses as they are inherent to, and incident upon a body, by the use of a model made from a photoelastic material. A polarised incident ray ol light is passed through a stressed model and is resolved into different rays that pass along the 'principal planes' acted upon by those stresses. According to the wave theory of light, the rays travel through the model at different velocities, as a temporary double refractive pattern proportional to the principal stresses in the photoelastic material. A photoelastic polariscope is used to measure the difference in velocities, and therefore the difference between, the principal stresses. Polarising filters are used to pass light waves parallel to the polarising axis. "Only those components of the incident wave parallel to the principal planes of the model at the incident surface will pass through, and these components are retarded in accordance with the degree of stress existing at the point in question", (Mahler & Peynton, 1955). The resultant light then passes through a second polarising filter oriented at 90 degrees to the first and a pattern emerges whereby light that has been altered in orientation by the principal planes will display interference phenomena. lf a white light is used then there will be the production of complementary colours to those wavelengths being retarded. The use of a monochromatic source will result in dark and light fringes which allow quantitative analysis. A polychromatic source produces coloured patterns that allow for additional 186 information. This improves the stress analysis by qualitative discrimination between compression and tension, (Baeten, 1975). There is a requirement to represent the original structure accurately to ensure vatid resutts, this being complicated by the wide diversity of biologic norms and lack of structural homogeneity, (Tanner, 1972; Caputo et al, 1974; McGuinness et at, 1992). Thus, morphology and environment for the model must simulate the actual body for valid interpretation of the bifringent refraction that occurs during analysis. To this end, it would seem appropriate to use materials that approximate the modulus of elasticity for the various biologic components being analysed, (Tanner, 1972). The modulus of elasticity for various components of the stomatognathic system are detailed in TABLE 6 - 1, where there is a cortical bone to periodontal ligament ratio o12.5 to 1, (Mehta et al, 1976). Young's modulus of elasticitY Williams et al,'84 Tanne et al,'87 Mehta et al,'76 N/mm2 N/mnf N/ mnÊ Enamel 8.41 x 104 Dentine 1.83 x 104 2.03 x 104 Pulp 2.03 PDL 0.05 - 9.99 x 103 6.9 x 103 1.3 x 104 CorticalBone 3.4 x 105 3.4 x 104 Cancellous Bone 1 .37 x 104 1.4 x 104 TABLE A. 1 Elastic (Young's) modulus for components of the stomatognathic sYstem. Adapted from McGuinness et al (1992) 187 A correlation between photoelastic models and histological section has been shown, adding credibility to the photoelastic technique, (Brodsky et al, 1975), given that the model design is accurate, (Mehta et al, 1975). Caution should be exercised in correlating a model to biological events, since no model can accurately simulate tooth movement and certain assumptions made may not hold for the 'in vivo' situation. These assumptions include that stresses are evenly distributed over the entire root surface, and that forces applied to the crown are those transmitted into, and by, the periodontium; both of which do not hold for the 'in vivo' environment, (McGuinness et â1, 1 992). The use of photoelasticity in relation to the study of dentistry was introduced by Zak (1935) who studied some of the alveolar effects of orthodontic mechanics, (cited by de Alba et al, 1979). Some studies since then have relevance to the current investigation into arch- form anchorage bend analysis, since they have looked at the effect of force level distribution from orthodontic appliances on the stomatognathic system. One theory of tooth movement relates the application of stress to the initiation of cellular events, (Sandy, 1992), and it would seem appropriate to use photoelastic studies to improve on the empirical understanding of force apptication in orthodontic therapy. This type of study should retain cognisance of "tissue response in relation to the direction, duration, continuity, and distance through which forces are applied" while allowing visualisation and analysis of the forces around the tooth root, (Caputo et al, 1974). Previous studies have investigated the application of photoelasticity to orthodontics. Nikolai & Schweiker, (1972\, found that tooth 188 displacement is proportional to the logarithm of applied force, but suggested that due to the hourglass shape of the incisal periodontal ligament, a linear relationship could be validated. Rodriguez & Arrechea, (1973)' found that a change in the direction of applied force will exhibit an alteration in the stress patterns around the root: axial forces result in an even distribution of the stress patterns, while lateral forces may increase the risk to pulp integrity because of excessive periapical stress concentration. Hayashi, et at, (1975), however, agree this is the case for a multi-rooted tooth, but for a single rooted tooth under lateral force, the major stress concentration is at the alveolar crest distal to the applied force direction, and tensite in nature. Thus, a correlation exists between tooth root morphology and stress distribution in the alveolus during application of orthodontic force. Atso, "lor a given movement (stress distribution is) independent of the force magnitude, but directly related to the direction of force application". The effects of second order, or vertical bends were studied by Caputo et al, (1974), when they considered the effect of gable bends for canine retraction using Blue ElgiloysT wire. Using o, 30, 45, and 60 degree bends, a relationship was noted between torque and translatory forces during bodily movement. Baeten, (1975), showed the complicated canine retraction mechanics of Ricketts and Burstone to be the least effective. A strong moment of simple design seems to be superior for producing bodily movement. Chaconas et al, (1989), suggest the type of malocclusion present should dictate the type of retraction used, particularly in relation to the depth of the bite on presentation. Stewart et al, (1978) found that Class 2 and Class 3 elastic forces on the teeth show a tendency for stress 37 Rocky Mountain Dental Products Company, Denver, Colorado 189 concentration both at the alveolar crest and the apex for the canines, and atong the mesial and apex of the mesial root of the molar. Photoelastic studies have also been applied to investigations into the effects of occlusal load. The anatomy and morphology of the masticatory apparatus seem to corretate with the type of occlusal loads applied, (Ward & Molnar, 1g8O). A change in angle of occlusal force application by 30 degrees results in a four-fold increase in compressive stress in the supporting bone mesial to the direction of force, (Hood et â1, 1975). Occlusal interferences, or premature contacts, show a tendency to concentrate particular stresses along one side of the tooth root, (Mehta et al' 1976). Studies have been undertaken to observe the effects of orthopædic traction on the facial skeleton. Photoelastic effects for various types of extraoral traction with different force levels support histologic and clinical observation in the use of those techniques as applied to both dental and cranio-facial structures, (de Alba et al, 1976; Chaconas et al, 1976; de Alba et al, 1979; Perez et al, 1980; de Alba et al, 1992; ltoh et al' 1985). Matsuura & Hanada, (1985) contend Finite Element Simulation allows for a superior systematic and quantitative analysis of orlhodontic tooth movement, since photoelastic fringe patterns correspond to differences in stresses rather than the actual principal stress. Nikolai & Schweiker, (1972), suggest that a sound correlation between elasticity analysis and the photoelastic model gives additional support to the theoretical model of tooth movement. Caputo et al, (1974) note that the patterns produced in photoelastic studies correlate well with those forces applied to achieve accepted observations of tooth movement.