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Assessment of Measurement Techniques to Determine the Interfacial Properties of Bilayer Dental Ceramics

Permanent Link: http://ufdc.ufl.edu/UFE0024307/00001

Material Information

Title: Assessment of Measurement Techniques to Determine the Interfacial Properties of Bilayer Dental Ceramics
Physical Description: 1 online resource (112 p.)
Language: english
Creator: Anunmana, Chuchai
Publisher: University of Florida
Place of Publication: Gainesville, Fla.
Publication Date: 2009

Subjects

Subjects / Keywords: bilayer, chevron, fractography, fracture, indentation, interfacial, mechanics, microtensile, notch, residual, stress, toughness, vickers, zirconia
Materials Science and Engineering -- Dissertations, Academic -- UF
Genre: Materials Science and Engineering thesis, Ph.D.
bibliography   ( marcgt )
theses   ( marcgt )
government publication (state, provincial, terriorial, dependent)   ( marcgt )
born-digital   ( sobekcm )
Electronic Thesis or Dissertation

Notes

Abstract: The clinical success of all-ceramic dental restorations depends on the quality of interfacial bonding between ceramic layers. In addition, the residual stress in the structure that developed during ceramic processing is one of the important factors that contributes to the quality of the bond. Because all-ceramic restorations are usually fabricated as bilayer or trilayer structures and failures of all-ceramic restorations have been frequently reported as chipping or delamination of the veneer layers, the interfacial quality of bilayer dental ceramic restorations was investigated. However, most of the published bond test data reflect strength values that are inversely related to cross-sectional areas and failure locations are frequently disregarded or bond strength values are misinterpreted. In addition, residual tensile stresses that develop in the structures because of thermal expansion/contraction mismatches may also adversely affect interfacial fracture resistance. The first objective of this study was to determine the interfacial toughness of bonded bilayer ceramics using two different approaches. The results indicate that the short-bar chevron-notch test and a controlled-flaw microtensile test can induce interfacial failure that represents true bonding quality. The second objective of this study was to test the hypothesis that residual stresses estimated from an indentation technique are not significantly different from residual stresses that are calculated based on fractography and flexural strength. The indentation technique may be useful as a simplified method to determine residual stresses in bilayer dental ceramics. The results of this study demonstrate that there is no significant difference in mean residual stresses determined from the two techniques. Because of relationship between residual stresses and apparent interfacial toughness, estimates of residual stresses can now be estimated more rapidly by measuring the apparent interfacial toughness of bilayer systems.
General Note: In the series University of Florida Digital Collections.
General Note: Includes vita.
Bibliography: Includes bibliographical references.
Source of Description: Description based on online resource; title from PDF title page.
Source of Description: This bibliographic record is available under the Creative Commons CC0 public domain dedication. The University of Florida Libraries, as creator of this bibliographic record, has waived all rights to it worldwide under copyright law, including all related and neighboring rights, to the extent allowed by law.
Statement of Responsibility: by Chuchai Anunmana.
Thesis: Thesis (Ph.D.)--University of Florida, 2009.
Local: Adviser: Anusavice, Kenneth J.
Local: Co-adviser: Mecholsky, John J.

Record Information

Source Institution: UFRGP
Rights Management: Applicable rights reserved.
Classification: lcc - LD1780 2009
System ID: UFE0024307:00001

Permanent Link: http://ufdc.ufl.edu/UFE0024307/00001

Material Information

Title: Assessment of Measurement Techniques to Determine the Interfacial Properties of Bilayer Dental Ceramics
Physical Description: 1 online resource (112 p.)
Language: english
Creator: Anunmana, Chuchai
Publisher: University of Florida
Place of Publication: Gainesville, Fla.
Publication Date: 2009

Subjects

Subjects / Keywords: bilayer, chevron, fractography, fracture, indentation, interfacial, mechanics, microtensile, notch, residual, stress, toughness, vickers, zirconia
Materials Science and Engineering -- Dissertations, Academic -- UF
Genre: Materials Science and Engineering thesis, Ph.D.
bibliography   ( marcgt )
theses   ( marcgt )
government publication (state, provincial, terriorial, dependent)   ( marcgt )
born-digital   ( sobekcm )
Electronic Thesis or Dissertation

Notes

Abstract: The clinical success of all-ceramic dental restorations depends on the quality of interfacial bonding between ceramic layers. In addition, the residual stress in the structure that developed during ceramic processing is one of the important factors that contributes to the quality of the bond. Because all-ceramic restorations are usually fabricated as bilayer or trilayer structures and failures of all-ceramic restorations have been frequently reported as chipping or delamination of the veneer layers, the interfacial quality of bilayer dental ceramic restorations was investigated. However, most of the published bond test data reflect strength values that are inversely related to cross-sectional areas and failure locations are frequently disregarded or bond strength values are misinterpreted. In addition, residual tensile stresses that develop in the structures because of thermal expansion/contraction mismatches may also adversely affect interfacial fracture resistance. The first objective of this study was to determine the interfacial toughness of bonded bilayer ceramics using two different approaches. The results indicate that the short-bar chevron-notch test and a controlled-flaw microtensile test can induce interfacial failure that represents true bonding quality. The second objective of this study was to test the hypothesis that residual stresses estimated from an indentation technique are not significantly different from residual stresses that are calculated based on fractography and flexural strength. The indentation technique may be useful as a simplified method to determine residual stresses in bilayer dental ceramics. The results of this study demonstrate that there is no significant difference in mean residual stresses determined from the two techniques. Because of relationship between residual stresses and apparent interfacial toughness, estimates of residual stresses can now be estimated more rapidly by measuring the apparent interfacial toughness of bilayer systems.
General Note: In the series University of Florida Digital Collections.
General Note: Includes vita.
Bibliography: Includes bibliographical references.
Source of Description: Description based on online resource; title from PDF title page.
Source of Description: This bibliographic record is available under the Creative Commons CC0 public domain dedication. The University of Florida Libraries, as creator of this bibliographic record, has waived all rights to it worldwide under copyright law, including all related and neighboring rights, to the extent allowed by law.
Statement of Responsibility: by Chuchai Anunmana.
Thesis: Thesis (Ph.D.)--University of Florida, 2009.
Local: Adviser: Anusavice, Kenneth J.
Local: Co-adviser: Mecholsky, John J.

Record Information

Source Institution: UFRGP
Rights Management: Applicable rights reserved.
Classification: lcc - LD1780 2009
System ID: UFE0024307:00001


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1 ASSESSMENT OF MEASUREMENT TECHNIQUES TO DETERMINE THE INTERFACIAL PROPERTIES OF BILAYER DENTAL CERAMICS By CHUCHAI ANUNMANA A DISSERTATION PRESENTED TO THE GRADUATE SCHOOL OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF DOCTOR OF PHILOSOPHY UNIVERSITY OF FLORIDA May 2009

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2 2009 C huchai Anunmana

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3 To my parents and my loving family, Siriwan, Pamarawee and Pete Anunmana

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4 ACKNOWLEDGMENTS First of all, I would like to thank Mahidol University, the Royal Thai Government and my country Thailand, for giving me an opportunity to have the highest education in one of the best educational place of the world. Also, I would like to thank all the faculty members at the Department of Pr osthodontics, Faculty of Dentistry for their encouragement and friendship. I would like to express my sincere gratitude to the staff in the Department of Dental Biomaterials who have assisted me through the Ph.D. program. Without their friendship, my life at the University of Florida would not have been so enjoyable and rewarding. To all my past and current fellow g raduate students for their help and friendship, I wish all of them success and great achievements from what they have learned My most sincere thanks go to my advisors, Dr. Kenneth J. Anusavice and Dr. John J. Mecholsky Jr. for their valuable advice and guidance during my graduate study They have taught me excellent lessons as a student and showed me how to be a good teacher They will always be my teachers and will always have my respect and friendship. My special thank s go to Mr. Robert Ben Lee who is always there for help. There is no doubt that I could not have finished m y research without him. I cannot describe in words how helpful and friendly a person he has been I wish Ben and his wife Rita all the happiness and good health. My very special thanks go to my parents and my wife Siriwan, and my daughter Pam. For my pare nts they have worked hard for me, taught me to be a good person and given me an education I will never be able to compensate them for their love and encouragement throughout my life. M y wife Siriwan sacrific ed he r career for my education. I deeply appreciate her dedication, support, patience and understanding that she has given me during my life. A final

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5 thanks and love to my daughter Pam who makes me the happiest person. I have never had a hard time in the USA because of her cheerfulness and free s pirit.

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6 TABLE OF CONTENTS ACKNOWLEDGMENTS ...............................................................................................................4 LIST OF TABLES ...........................................................................................................................8 LIST OF FIGURES .........................................................................................................................9 ABSTRACT ...................................................................................................................................12 1 INTRODUCTION ..................................................................................................................14 2 FRACTURE TOUGHNESS OF CERAMIC CORES AND GLASS VENEERS .................20 2.1 Materials and Methods .....................................................................................................21 2.1.1 Specimen Fabrication .............................................................................................21 2.1.2 Fracture Toughness Testing ...................................................................................22 2.1.3 X ray Diffraction Analysis (XRD) .........................................................................24 2.1.4 Scanning Electron M icroscopy (SEM) ...................................................................25 2.1.5 Statistical Analysis .................................................................................................25 2.2 Re sults ...............................................................................................................................25 2.3 Discussion .........................................................................................................................27 3 INTERFACIAL TOUGHNESS OF BILAYER DENTAL CERAMICS BASED ON A SHORT BAR, CHEVRONNOTCH TEST ...........................................................................37 3.1 Materi als and Methods .....................................................................................................39 3.1.1 Material Combinations ...........................................................................................39 3.1.2 Sample Preparations ...............................................................................................39 3.2 Results ...............................................................................................................................41 3.3 Discussion .........................................................................................................................42 3.4 Conclusions .......................................................................................................................43 4 INTERFACIAL TOUGHNESS USING A CONTROLLEDFLAW MICROTENSILE TEST .......................................................................................................................................47 4.1 Materials and Methods .....................................................................................................48 4.1.1 Material Combinations ...........................................................................................48 4.1.2 Specimen Preparation .............................................................................................48 4.2 Results ...............................................................................................................................50 4.3 Discussion .........................................................................................................................51 4.4 Conclusions .......................................................................................................................52 5 EFFECT OF CROSS SECTIONAL AREA ON INTERFACIAL BOND STRENGTH AND INTERFACIAL TOUGHNESS OF BILAYER DENTAL CERAMICS .....................57 5.1 Materials and Methods .....................................................................................................58

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7 5.1.1 Specimen Prepar ation .............................................................................................58 5.1.2 Interfacial Toughness Determination .....................................................................59 5.1.3 Microstructural Characterization of Interfacial Zones ...........................................59 5.2 Results ...............................................................................................................................60 5.3 Discussion .........................................................................................................................61 6 RESIDUAL STRESS IN GLASS: INDENTATION CRACK AND FRACTOGRAPHY APPROACHES ......................................................................................................................71 6.1 Materials and Methods .....................................................................................................74 6.1.1 Specimen Preparation .............................................................................................74 6.1.2 Fracture Toughness Test .........................................................................................74 6.1.3 Statistical Analysis .................................................................................................75 6.2 Results ...............................................................................................................................76 6.3 Discussion .........................................................................................................................76 7 RESIDUAL STRESSES IN BILAYER ALL CERAMIC SYSTEMS ..................................82 7.1 Materials and Methods .....................................................................................................83 7.1.1 Specimen Preparation .............................................................................................83 7.1.1.1 Monolithic bars cont rol groups ....................................................................83 7.1.1.2 Bilayer bar specimens ..................................................................................84 7.1.2 Testing Methods .....................................................................................................84 7.1.3 Statistical Analysis .................................................................................................86 7.2 Results ...............................................................................................................................87 7.3 Discussion .........................................................................................................................88 7.3 Conclusions .......................................................................................................................90 8 CONCLUSIONS ....................................................................................................................94 APPENDIX ....................................................................................................................................98 A FLEXURAL STRENGTH CALCULATED FROM COMPOSITE BEAMS .......................98 LIST OF REFERENCES .............................................................................................................103 BIOGRAPHICAL SKETCH .......................................................................................................112

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8 LIST OF TABLES Table page 21 Dental ceramics, compositions, firing temperature, and annealing temperature ...............29 22 Dental ceramics and programmed firing parameters .........................................................29 23 Elastic modulus, hardness, indent ation load, indentation strength (MPa) and fracture toughness [KIC] (MPam1/2) calculated from fractography (F) ..........................................30 31 Materials, ceramic type s and processing temperature s .....................................................44 32 Mean failure load and the apparent t oughness ...................................................................44 41 Materials, types, and processing temperature ....................................................................52 42 Mean interfacial toughness ( SD) and statistical subsets .................................................52 51 Mean interfacial toughness and mean bond strength (SD) of ceramic groups with varying cross sectional area. ..............................................................................................64 61 Fracture strength (MPa) and fracture toughness (MPa m1/2) of soda lime glass ...............78 62 Residual stress (MPa) of each specimen based on the indentation technique and the superposition of the stress intensity factor method ............................................................78 71 f), crack length, apparent fracture toughness (Kc), and R) ..........................................................................................................91 72 R (MPa) of each specimen using the indentation technique and fractography .......................................................................................................................91 81 Interfacial toughness of bilayer specimens ........................................................................97

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9 LIST OF FIGURES Figure page 21 Illustration of the determination of the glass transition temperature (Tg) of the LC group (IPS e.max Press) from the dilatometric curve.. ...................................................30 22 Schematic of a brittle material illustrating the critical flaw siz e [25] ................................31 23 Optical microscopy images of fracture surfaces of bar specimens : (a) yttria stabilized zirconia core ceramic (ZC); (b) lithia dis ilicate glass ceramic core (LC) ........................31 24 X ray diffraction pattern of an e.max Ceram veneering glass specimen (GV) ..............32 25 X ray diffraction pattern of an e.max ZirPress veneering glass specimen (HV ) ............32 26 X ray diffraction pattern of an e.max Press core ceramic specimen (LC) ......................32 27 X ray diffraction pattern of an as sintered e.max ZirCAD core ceramic specimen (ZC) ....................................................................................................................................33 28 X ray diffraction pattern of a ZC core ceramic specimen after polishing. An arrow indicates the monoclinic peak of zirconia. .........................................................................33 29 X ray diffraction pattern of a ZC specimen after grit blasting by 100 m alumina for 1 min ..................................................................................................................................34 210 X ray diffraction patterns: (a) sandblasted by 100 m alumina; (b) sandblasted after heat treatment at 1050 C for 15 min (t = tetragonal phase, m = monoclinic phase) ........34 211 SEM micrographs of e.max Ceram (GV) etched with an aqueous solution of 30% H2SO4 and 4% HF for 10 s at (a) 500x, and (b) 30000x ....................................................35 212 SEM micrographs of e.max ZirPress (HV) etched with an aqueous solution of 30% H2SO4 and 4% HF for 10 s; (a) 5000x; (b) 20000x ...........................................................35 213 SEM micrograph (5000x) of e.max Press (LC) etched with an aqueous solution of 30% H2SO4 and 4% HF for 10 s ........................................................................................36 214 SEM micrograph (30000x) of e.max ZirCAD (ZC) thermally etched at 1420 C for 15 min ................................................................................................................................36 31 Specimen geometry: (a) A short bar geometry was modified for two dissimilar materials; (b) Schematic of crack plane of a specimen ......................................................44 32 Specimen preparation: (a) Core ceramic before veneering; (b) V shaped notch preparation; (c) Completed test specimen; (d) Specimen in the test fixture. .....................45 33 A fractured specimen of the LC/GV group ........................................................................45

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10 34 Load displacement curve of a chevron notch, short bar specimen ...................................46 35 (a) SEM image of fracture surface of chevron notch specim en of the LC/GV group, and (b) O ptical microscopy image .....................................................................................46 41 Specimen preparation .........................................................................................................53 42 Inte rfacial area of an LC/GV specimen .............................................................................53 43 An indent at the interface between core ceramic and veneer ceramic ...............................54 44 Specimen attached to the loading grips .............................................................................54 45 Optical micrograph of a critical flaw .................................................................................55 46 SEM micrograph of the critical flaw shown in Figure 45 ................................................55 47 Box plot of interfacial toughness determined from a controlled flaw microtensile test ....56 51 SEM micrograph of a fracture surface of a microtensile bar in the ZC/CS group. White arrows indicate a critical crack boundary ................................................................65 52 Energy dispersive (X ray) spectrum of e.max Ceram veneering ceramic (CS) .............65 53 Energy dispersive (X ray) spectrum of e.max ZirCAD core ceramic (ZC) ...................66 54 Energy dispersive (X ray) spectrum of ZirPress veneering ceramic (HIP) ....................66 55 X ray intensity of elements across the interface of the ZC/CS specimens from EPMA using the line scan mode ....................................................................................................67 56 X ray intensity of elements across the interface of the ZC/HIP specimens from EPMA using the line scan mode ........................................................................................68 57 X ray intensity of elements across the interface of the LC/CS specimens from EPMA using the line scan mode ....................................................................................................69 58 SEM micrographs of a microtensile bar from (a) the CS group (2.25 mm2), and (b) the HIP group (1.44 mm2) showing distribution of porosity .............................................70 59 Polished surface from (a) a HIP glass veneer specimen, and (b) a CS glass veneer specimen showing differences in the size and distribution of pores ..................................70 61 Schematic of the design of heat strengthening apparatus [Adapted from Anusavice et.al ., 1989 [109]] ...............................................................................................................79 62 Schematic of bar specimen showing an indentation crack used for residual stress calculation. .........................................................................................................................79

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11 63 Optical microscopy images of the fracture surface showing a critical crack of a fracture toughness test specimen .......................................................................................80 64 Optical microscopy images of an indentation crack induced at a load of 19.6 N in (a) an annealed specimen and (b) a specimen with compressive residual stress. ...................80 65 Schematic of a semicircular crack with indentation crack depth [a] within the global compressive residual stress region on the side R] under uniform flexural A] [107]. ..........................................................................................81 71 Schematic of specimen dimensions of bilayer specimens showing indentation crack used for calculation of residual stress. ...............................................................................92 72 A fracture surface of a bilayer specimen of the LC/GV group. Arrow demonstrates the critical flaw from an original indent. ...........................................................................92 73 Fracture surface of a specimen from the ZC/GV group showing a critical crack boundary (black arrows) and the original indent (white arrow). .......................................93 74 Critical flaw on the fracture surface of a b ilayer specimen from the LC/GV group (a) optical micrograph and (b) SEM micrograph. ...................................................................93 A 1 Cross section schematic of the bilayer b eam of the LC and GV. ....................................102 A 2 Transformed section schematic of the bilayer beam of the LC and GV .........................102

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12 ABSTRACT OF DISSERTATION PRESENTED TO THE GRADUATE SCHOOL OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF DOCTOR OF PHILOSOPHY ASSESSMENT OF TECHNIQUES TO DETERMINE THE INTERFACIAL PROPERTIES OF BILAYER DENTAL CERAMICS By Chuchai Anunmana May 2009 Chair: Kenneth J. Anusavice Cochair: John J. Mecholsky Major: Materials Science and Engineering The clinical success of all ceramic dental restorations depends on the quality of interfacial bonding between ceramic layers In addition, the residual stress in the structure that developed during ceramic processing is one of the important factors that contributes to the quality of the bond. Because all ceramic restoration s are usually fabricated as bilayer or trilayer structures and failures of all ceramic restorations have been frequently reported as chipping or delamination of the veneer layers, the i nterfacial quality of bilayer dental ceramic restorations was investigated. However, most of the published bond test data reflect strength values that are inversely related to cross sectional areas and failure locations are frequently disregarded or bond s trength values are misinterpreted In addition, residual tensile stresses that develop in the structures because of thermal expansion/contraction mismatches may also adversely affect interfacial fracture resistance. The first objective of this study was to determine the interfacial toughness of bonded bilayer ceramics using two different approaches. The result s indicate that the short bar chevronnotch test and a controlledflaw microtensile test can induce interfacial failure that represents true bonding quality The second objective of this study was to test the hypothesis that residual stresses estimated from an indentation technique are not significantly different from residual

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13 stresses that are calculated based on fractography and flexural strength T he indentation technique may be useful as a simplified method to determine residual stresses in bilayer dental ceramics. The results of this study demonstrate that there is no significant difference in mean residual stresses determined from the two techniques. Because of relationship between residual stresses and apparent interfacial toughness, estimates of residual stresses can now be estimated more rapidly by measuring the apparent interfacial toughness of bilayer systems

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14 CHAPTER 1 INTRODUCTION The demand for high quality, durable a llceramic restorations has increased over the past decade because of their esthetic capability and biocompatibility Recently, new dental core ceramics with high strength and high fracture toughness were introduced as alternatives to metals However, relatively high failure rates of some ceramic systems [1] raise serious concerns on the benefits of using all ceramic prostheses before the timedependent probability of structure d components is adequately established. Dental ceramics consist primarily of glasses, feldspathic bonded ceramics, glass ceramics or highly crystalline ceramics and can be classified by their composition, processing methods, microstructure, firing tempera tures, and translucency [2] Highly esthetic ceramics, such as veneering cer amics, are predominantly composed of glass pha s e s and they best simulate the optical properties of enamel and dentin. Core ceramics with greater flexural strength and toughness are generally more opaque and crystalline [3] Naturally occurring mineral components of veneering ceramics are composed primarily of potash (K2O) and soda (Na2O). They also contain silica (SiO2) and alumina (Al2O3) [1, 2] Glasses based on feldspar used in dental opaque, gingival, and incisal ceramics are resistant to crystallization during firing and they have a broad firing range. Ceramics based on K2O Na2O SiO2Al2O3 systems contain 1115% K2O. T hey can form leucite (K2O Al2O34SiO2) as the principal crystalline phase [4] When leucite is added to feldspathic glasses it creates thermally compatible dental porcelains that bond chemically to dental alloys to produce metal ceramic prostheses [1 3] Technically, the term porcelain is incorrect since dental ceramics do not contain clay as a component.

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15 Glass ceramics are polycrystalline materials, whose crystallization by nucleation and crystal growth are controlled by special heat treatments [5, 6] Because of the small amount of glass phase, these materials tend to be relatively porefree. Dicor (Dentsply) was the first glass ceramic used in dentistry. It contained about 55% tetrasilicic fluoromica by weight KMg2. 5Si4O10F2 [7] However, Dicor has been discontinued because of its low tensile strength and the need for external coloring. Nevertheless, in one study, 76% of Dicor crowns survived after 14 years if their fitting surfaces had been acid etched compared with a survival of 50% for a nonetched group [1] Recently, two glass ceramics containing 70 vol% of elongated lithia disilicate crystals have been introduced as core ceramics (Empress 2 and e.max Press, Ivoclar Vivadent, Schaan, Liechtenstein). Ceramic restorations that are made from glass ceramics are more translucent than purely crystalline ceramics because of their glassy phases. However, because of their low fracture toughness these systems are limited to three unit fixed partial dentures that replace a missing tooth anterior to the second premolar [8] Polycrystalline dental ceramics such as alumina and zirconia have no glassy phase, and they are relatively opaque compared with ceramics that contain a glass phase. Therefore, these stronger and tougher material s are used as substructures (cores) to support weaker but more translucent glass veneers. The ceramics in these systems include a densely sintered high purity alumina and partially stabilized zirconia that can be milled to dense anatomic shapes by a CADCAM devic e in either partially sintered or fully dense form. The ceramic systems based on zirconia have excellent mechanical properties compared with the other ceramic systems Therefore, they have been used as substructures with a substantial reductio n in core thickness to support the glass veneer in posterior fixed partial denture s Pure z irconia has three allotropes (cubic, tetragonal, and monoclinic) depending on the

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16 temper ature at which they are pressed [9, 10] At ambient c onditions and upon heating up to 1170 C zirconia exists in the monoclinic (m) phase The structure is tetragonal (t) between 1170 C to 2370 C and cubic (c) above 2370 C [10, 11] The transformation from tetragonal to monoclinic whic h occurs at approximately 1000 C, causes cracking because of a 34 % volume expansion [10, 11] This cracking can result in catastrophic fracture of the fabricated components. Blending zirconia with small amount s of stabilizing oxides, such as yttria (Y2O3), ceria (CeO2), calcia ( CaO), or magnesia (MgO) maintain the metastable tetragonal phase at room temperature [11] Fracture toughness of partially stabilized zirconia can be enhanced by inducing the m transformat ion in the stress field of a propagating crack [12] This process efficiently arrests crack propagation and lead s to increased fracture toughness. Dental ceramic restorations based on zirconia systems have gain ed in popularity and interest among dentists and researchers because of their excellent mechanical properties compared with other ceramic systems. Their mechanical properties are the great est ever reported for any core ceramics; however, they still need to be veneered using glassy phase ceramics. These high toughness ceramic systems are usually fabricat ed using CAD CAM technology, and they allow thinner core layers (~0.5 mm) to be produced that can be used for construction of all ceramic fixed partial dentures in posterior regions [13] The commercial examples of these materials include Cercon (Dentsply), DC Zirkon (DCS), IPS e.max ZirCAD (Ivocl ar Vivadent, Schaan, Li ech t enstein ), LAVA Frame (3M ESPE), and Nanozir (Matsushita Electric Works, Osaka, Japan) Currently, the application of high strength and hi gh toughness core materials has allowed all ceramic restorations to be employed in posterior locations. However, in comparison to metal ceramic restorations which are considered as a gold standard for development of new materials,

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17 a 2007 study indicated that the survival rate of all ceramic FPDs was significantly less than those of metal ceramic FPDs [14] In addi tion, the frequencies of framework and veneering ceramic fractures were significantly greater for all ceramic FPDs [14] However, fractures of zirconia frameworks have been rarely reported. Failures of all ceramic restorations have been most often reported as chi pping of the glass veneer layer [1422] Chipping of glass veneer has been reported as early as the first year of service [15] I n one case, 15% of the restorations chipped within two years [22] These failures may be caused by the weak bonding between ceramic cores and g lass veneers and/ or by the residual stresses developed during processing. Since the interface plays an important role in the mechanical performance of d ental ceramic restorations that are usually fabricated as bilayer structure s, it is surprising that very few studies have been performed to determine the interfacial fracture resistance of these all ceramic bilayer structures. Additionally, no standard test has been proposed to measure interfacial bonding or interfacial toughness of dental all ceramics composites. Shear and microtensile bond test s have been used to test the bond strength between bonded bilayer ceramic structures [23] However, the commonly used shear bond test often produces fracture at a distance from the interfaces [24] which is caused by a tensile stress component of bending and a nonuniform stress distribution [18, 19] F or the microtensile bond test, bond strength varies with the specimen shape and dimensions [25] Critical f law size from fractographic analysis and calculated fracture stress are used to measure the critical stress intensity factor ( KIC) of monolithic materials [26, 27] When a crack is initiated along the interfacial zone, the apparent interfacial toughness or bonding integrity is estimated from crack size measurements on fracture surfaces and fracture strength using fracture

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18 mechanics principles This toughness re presents the ability of the interface to resist unstable crack propagation [28, 29] Residual stresses are usually present in bilayer dental ceramics because of differences in the coefficient s of thermal expansion/contraction between the core and veneer ceramics [2] In metal ceramic systems, it is generally thought that a slightly great er coefficient of thermal contraction for the core ceramic com pared with the glass veneer (positive mismatch) is beneficial because compressive residual stresses can be developed across the metal/veneer interface in the weaker ceramic veneer layers. However, the differences in thermal expansion or contraction coefficients sh ould not be too great since catastrophic radial tensile stresses may also develop. Moreover, it is uncertain whether or not the positive mismatch is advantageous for all ceramic systems. An i ndentation technique that is based on a pyramidal diamond indenter has been widely used to measure mechanical properties such as fracture toughness, hardness and a brittleness index [30] This technique has also been applied to determine residual stresses [31] However, the indentation technique has been criticized as an unreliable method to determine any fracture resistance parameter [32] The first objective of this study was to determine the apparent interfacial toughness between bonded bilayer dental ceramics using the following two different approaches: (1) chev r onnotch short bar specimens were used to produce controlled failures along the interfacial zone [chapter 3] ; (2) a control led flaw microtensile test wa s employed in conjunction with fractography for measuri ng the interfacial toughness [chapter 4] Additionally, the assumption that a control led flaw microtensile test for analysis of interfacial toughness is not dependent on specimen dimensions in contrast to a traditional bond stre ngth test was examined [chap ter 5]

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19 The second objective of the study was to verify that the indentation crack tec hnique can be used as a simplified method for determining residual stress es i n soda lime silica glass [chapter 6] This technique was compared with the method that requir es flexural strength combined with a fractographic analysis parameter s In the second part of this study, both techniques were appli ed to determine the residual surface stresses in bilayer dental ceramics [chapter 7] The results are summarized and the mai n conclusions of this work are stated in final chapter.

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20 CHAPTER 2 FRACTURE TOUGHNESS OF CERAMIC CORES AND GLASS VENEERS Fracture toughness represents a material s resistance to rapid crack propagation and it can be characterized by the parameter, KIC [33] KIC is an intrinsic property which, in contrast to strength, is generally not dependent on the size of the initiating cracks. Strength is sometimes used as an indicator for material performance of brittle materials. Howe ver, strength alone cannot provide sufficient information to predict structural performance [33, 34] Because strength is sensitive to specimen size, defects or the direction of surface grinding [35] strength is not considered as an intrinsic material property. Fracture surface analysis by fractography has been used as a quantitative tool to determine the toughness, origin of the crack initiation site the stress state and stress level at failure. Fractography provides a quantification of the stress level and the toughness of a material when one parameter is known once the location and size of the origin is determined The i ndentation technique has been used to investig ate mechanical properties such as hardness, toughness [36] and residual stress. It also provides a rapid evaluation of the mechanical properties of materials. When an indentation crack is introduced as a controlled flaw and is followed by a flexure test known as the indentation strength technique, fracture toughness can be determined from fracture str ength alone without fractographic information [37, 38] or it can be determined from fracture strength and critical crack size using fractography [26, 39] Veneering ceramics are composed predominantly of a glass phase T herefore, they tend to simulate the optical properties of enamel and dentin. Glass ceramics consist of a glass phase and one or more crystalline phases whose nucleation and crystallization are controlled by heat treatment [40, 41] Glass ceramics, by def inition, contain more than 50 % of crystalline phase by volume T ypically industrial glass ceramics contain more than 90% of a crystalline phase [42]

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21 Glass ceramics have greater strength and toughness than glassy phase ceramics because they are more crystalline, and they can be highly translucent and esthetic [40] Fully c rystalline ceramics such as alumina or zirconia have little or no glassy phase and have the greatest strength and toughness among all dental ceramics. They are relatively opaque compared with ceramics that contain a glass matrix. T herefore they are used as substructures or cores to support weaker but more translucent glass veneers. The objective of this chapter is to measure the fracture toughness of four dental ceramics used in this study. The results are necessary for further studies to calculate the apparent interfacial to ughness and residual stress distribution in bilayer dental ceramics. In addition, this study compared fracture toughness derived from the in dentation strength technique and to that derived from the fractography approach. 2.1 Materials and Methods 2.1.1 Specimen Fabrication The d ental ceramics, firing temperature, ceramic types and annealing temperatures employed in this study are presented in Table 2 1. P rogrammed firing conditions are shown in T able 2 2. For the glass veneer group [GV], IPS e.max Ceram (Ivoclar Vivadent AG, Schaan, Liechtenstein) powders were mixed as a slurry and condensed in a polyoxymethylene plastic mold and fired accor ding to manufact urers recommendations (Table 2 2 ). Yttria stabilized zirconia core ceramic bars (IPS e.max ZirCAD Ivoclar Viv adent AG, Schaan, Liechtenstein ) [ZC] were obtained from a presintered ceramic block. The presintered block was milled to oversi zed bar dimensions and sintered by the manufacturer The hot pressed fluorapatite glass veneer (IPS e.max ZirPress Ivoclar Viv adent AG, Schaan, Liechtenstein ) and hot press ed lithia disilicate glass ceramic core (IPS e.max Press Ivoclar Viv adent AG, Schaan, Liechtenstein ) were obtained from resin patterns using the lost wax technique. Resin patterns

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22 were invested in a silicone mold. The invested patterns were burn ed out for 60 min at 850 C in a preheated furnace (Radiance, Jelrus Int., Hicksville, USA) The mold cavity was filled by hot pressing ceramic ingots using a pressing furnace (EP 500, Ivoclar Vivadent AG, Schaan, Liechtenstein) at 930 C. The cooled specimens were divested and the interaction layer was removed from the specimen surface by grit blasting with 80 m glass beads at a pressure of 0.28 MPa. The bar specimens were cleaned in 1% HF for 30 min to remove the remaining investment. The L C specimens were grit blasted with 100 m Al2O3 particles at a pressure of 0.1 MPa for 30 s All specimens were polished initially with a 45 m abrasive by a metallographic polisher and sequentially with finer abrasive size, terminated with 1 m alumina abrasive (Mark V Laboratory, East Granby, CT, U.S.A.). All edge s were rounded to minimize str ess concentration. Final specimen dimension s of 4.5 mm x 2 mm x 28 mm were obtained after polishing. 2.1.2 Fracture Toughness Testing All bar specimens were annealed at approximately 50 C above the glass transition temperature (Tg) for 30 min (Table 21) and cooled slowly to eliminate all residual stresses. At this annealing temperature residual stresses we re eliminated in the matter of minutes [43] The Tg of the GV and the HV groups were reported by the manufacturer (Scientific documentation, Ivoclar Vivadent). To determine the Tg of the LC group, bar specimens with dimensions of 5 mm x 5 mm x 25.4 mm were prepared and the Tg was obtained using a dilatometer (Orton Dilatometer, Orton Ceramic Foundation, Colum bus, OH) at a heating/cooling rate of 5 K/min. Tg was determined from the intersection of two tangents to the straight line portions of the corresponding 0 versus temperature plot above and below the inflection point that corresponds to the transition from the vitreous state to metastable liquid phase of the material [44, 45] In this study, the contraction curve was used because of the release of trapped excess volume and creep during the heating curve. T herefore, the contraction curve is considered to be

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23 more reliable [46] Annealed specimens were slowl y cooled in the oven to room temper ature in 30 min. For the ZC group, the specimens were re transform ed at 1050 C in order to reverse changes in the monoclinic phase back to the stabilized tetragonal phase. The transformation from tetragonal to monoclinic phase may occur during grinding, fi nishing and polishing processes. For the GV and HV group, indentation cracks were induced at the center of polished specimens (6 per group) using a Vickers indenter at a load of 9.8 N. For the HC and ZC group, indentation loads of 19.6 N and 196 N were use d, respectively. These indent loads were selected such that the cracks were formed from the ti p of the Vickers indenter without excessive cracking or chipping. Crack lengths were approximately 2 3 times longer than the size of the diamond indenter. Specime ns were subje cted to four point flexure with a 20.0 mm lower span and a 6.7 mm upper span using an Instron universal testing instrument (Model 5500R) at a crosshead speed of 0.5 mm/min until failure occurred. Failure loads for the bar specimens were obtai ned, and the flexural f) were calculated using equation: f P L bh2 (2 1) where P = fracture load; L = test span; b = specimen width; and h = specimen thickness. Fracture toughness was calculated using the indentation strength technique and equation 22: KC v RE H 1 8 fP1 3 3 4 (2 2) where v R is a calibration constant (~ 0.59) In this case, the uncertainty in E/H is relatively unimportant and replacement of v RE H 1 8 by an average quantity ( 0.88) is reasonable. This

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24 value would fall within 10% of the error range in the Kc evaluation whose elastic/plastic parameters are totally unknown [37] In addition, fracture surfaces of broken bars were analyzed for the critical crack size using fractography and fracture toughness was calculated us ing equation 23: KIc f c1/2 (2 3) where KI c = fracture toughness; Y = numerical constant that accounts for loading and crack geometry (1.65 for indentation cracks); and c = radius of an equivalent semicircular crack (equal to [ab]1/2) [26] Figure 22 shows the fracture surface features that usually occur in brittle materials [26] To determine the fracture origins of the spe cimens, an optical microscope (OmniMet Modular Imaging System, Buehler, Lake Bluff, IL) was used. The fractured specimens in the LC group and the ZC group were sputter coated with goldpalladium to facilitate detection of critical cracks. 2.1.3 X ray Diffr action Analysi s (XRD) X ray diffraction was performed to determine the crystal phases in glass veneers and ceramic cores using a diffractometer ( Phillips APD 3720 XRD) Disk specimens with a diameter of 10 mm and a thickness of 1 mm were prepared for each material. For the zirconia core, the cryst al phases were determined for as received specimen s and separate specimens were prepared for grind ing, sandblasting, and re transformatio n procedures A grou nd specimen was handpolished with 45m grit metallograph ic polisher at 5000 rpm for 1 min. The sandblasted specimen was grit blasted with 100 m alumina at 0.5 MPa for 1 min. For the regeneration process, the ground and grit blasted specimens were heat treated at 1050 C for 15 min. The zirconia specimens were analyzed for possible phase changes induced by grind ing, sandblasting and regeneration processes. The diffraction patterns were compared to the patterns for standards

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25 from software that are based on the ICDD reference (Internat ional Center of Diffraction Data ) [ P CPDFWIN version 2.3, JCPDS ICDD] 2.1.4 Scanning Electron M icroscopy (SEM) Microstructural features were observed by SEM. Specimens were embedded in epoxy resin and polished initially with 45 m alumina abrasive and carr ied through to 0.05 m alumina abrasive Because of the differences in composition of each ceramic, the specimens were treated as follows : Yttria stabilized zirconia core specimen s (ZC) were thermally etched at 1420 C for 15 min Lithia disilicate glass c eramic core specimen s (LC) were etched with 4% aqueous solution of hydrofluoric acid (HF) for 10 s Fluorapatite condensed slurry glass veneer (GV) and f luorapatite hot pressed glass veneer (HV) were etched with an aqueous solution of 30% sulfuric acid (H2S O4) and 4% hydrofluoric acid (HF) for 10 s All specimens were carbon coated before SEM analysis. 2.1.5 Statistical Analysis The difference s between the mean fracture toughness by group was analyzed using a oneway analysis of variance and post hoc comparison using Tukey tests at a significance 0.05 (SAS for Windows 9.1.3 service pack 4, SAS institute, Cary, NC, USA.). Additionally, the toughness calculated from indentation s trength and that determined from the strength / fractography method were compared using a pair ed t 2.2 Results Figure 2 3 shows the fracture surfaces and the critical cracks for the two core ceramics and two veneer ceramics. All specimens fr actured from the indentation cracks. The means and standard deviations of indentation strength and the toughness of dental ceramics determined from both technique s a re summarized in T able 2 3. One way analysis of variance revealed significant differences a mong the mean values of fracture toughness f or the fou r ceramics used in this study

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26 (p < 0.0001). The mean KIC of the ZC ceramic was great est among all groups, and the mean KIC of the LC ceramic was significantly grea ter than that for the GV and the HV groups (p < 0.0001). There was no significant difference between the means of GV and HV group values (p > 0.05). A t test revealed that the mean toughness calculated from the indentation strength technique was significantly greater than the values obtaine d from fractography (p < 0.05) except for the LC group (p > 0.05). X ray diffraction analyses of each ceramic a re presented in F igure s 2 4 to 27. N o crystalline phase was identifiable in the glass veneers although it was shown by the manufacturer that flu orapatite nano crystals were present in glass veneer. The XRD pattern demonstrated that there were lithia disilicate crystal peaks (ICDD pattern # 400376) which were most prominent at 2 = 23.9, 24.4 and 24.9 in the LC specimen (Figure 2 6). For the zirconia specimens, the results from XRD showed there was an additional phase after grind ing and sandblasting procedures compared with the original tetragonal structure peak (ICDD pattern # 830113) The monoclinic phase was identifi ed within the tetragonal phase at 2 = 28.2 (ICDD pattern # 830944) which was the great est intensity peak for monoclinic zirconia for ground and sandblasted specimens (Figures 28 and 29). However, after heat treatment, there was no monoclinic phase det ected in the specimens (Figure 2 10). Microstructures of the ceramic materials used in this study a re presented in F igure s 211 to 214. For the glass veneer specimens (GV and HV), nanometer sized crystals were found and appeared to be the fluorapatite pha se as previously identified by the manufacturer (Figure 2 11 and 212) However, because of the limitation of X ray instrument s the structure of these crystals could not be identified by either the EDS or XRD methods used in this study. N eedle shaped crys tals were found in the LC specimen and those crystals were identified as lithia -

PAGE 27

27 disilicate (Li2Si2O5) crystals (Figure 2 13) Figure 214 shows densely packed submicron zirconia grains. 2.3 Discussion As expected, yttriastabilize d zirconia, which is a polycrystalline ceramic with no glass phase, has the greates t fracture toughness among all groups. The potential stressinduced phase transformation involves the transformation of metastable tetragonal phase to the monoclinic phase at the crack tip. This is accompanied by a volume expansion which induces localized compressive stress as the generalized tensile stress increases and crack propagation is impeded or arrested [9, 47] The mechanical properties of glass ceramics are usually superior to those of the same glass from which they are formed [48, 49] and the effect of crystallization on mechanic al properties of glass ceramics has been previously studied [48, 5052] The increase in toughness of lithia disilicate glass ceramics was attributed to the increase in the nucleation treatment and crystal volume fraction. The fracture toughness calculated from the indentation strength technique (IS) was significantly greater than that from fracture strength/ fractography method (F) in all groups except the LC group, for which the difference between means were not statistically significant The mean toughness derived from IS technique was not always comparable to that determined from other standard methods [38] Moreover, the strength value used to calculate t he toughness depends on loading rate because of the effect of subcritical crack growth which normally occurs in dental ceramics [53, 54] The greater the loading rate, the greater is the strength. In this study, the crosshead speed at 0.5 mm/min was used for the strength test. The stressing rate can be calculated from equation (2 4): E (2 4)

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28 where = stres s rate; E = elastic modulus; and = strain rate. The strain rate for each specimen was determined by modifying the crosshead displacement calculation using following ASTM Standard 079096 for a load span equal to one third of the support span [55] : D 0.21 L2d (2 5) where D = midsp an deflection rate (crosshead speed), = the specimen strain rate, L = the support span, and d = the depth of the beam. For the crosshead speed used in this study, the calculated stress rates of the ZC, LC, and GV an d HV we re 41, 19 and 13 MPa/s respectively. In contrast to the strength test, which is use d for calculating the toughness in the IS technique, the calculated toughness from a controlledflaw strength test and fractography is not dependent on the stress ing rate. A r ecent study [53] reported the effect of the testing environment and stressin g rate on the flexural strength and critical flaw size It was found that the specimens tested in an inert environment (in oil and high stressing rate) had significantly great er strength and smaller flaw size than those tested in water, but they did not exhibit a significant difference in fracture toughness [56] In this study, the same technique was used to investigate fracture toughness. In this study, XRD was used to determine the crystal phase in ceramic specimens. The diffraction patterns from the glass veneers were primarily amorphous because no sharp peaks were found in either of the two glass veneers. However, the SEM images show that there were crystal phases formed in the se materials (Figure s 211 and 212) [57] XRD is most sensitive to hi gh atomic weight elements, since the diffracted peak intensity from high atomic weight elements are much greater than from low atomic weight elements [58] Additionally there was only a small amount of fluorine (< 1%) detected in the glass veneer using electron probe microanalysis which was close to the fluorine content reported by the manufacturer. These are the most likely reason s why the crystal phases could not be detected by XRD.

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29 The comparative effect of surface grinding and sandblasting on yttria stabilized zirconia (Y TZP) and ceriastabilized zirconia /alumina (Ce TZP /Al2O3) has been studied previously [5961] The transformations of tetragonal (t) to monoclinic (m) phase were observed from surface grinding and sandblasting using XRD [5961] and Raman spectroscopy [60] Biaxial flexural strength was increased as a result of sandblasting the surface and the t more pronounced from sandblasting than from surface grinding [59] However, this effect may not be favorable when glass veneers are applied to the surface of core substructures because this might change the thermal contraction mismatch or the bonding ability of glass veneer to the ceramic cores. Table 2 1. Dental ceramics, compositions, firing temperature, and annealing temperature Dental Ceramics Composition Firing Temperature (C) Annealing temperature (C) IPS e.max Ceram (GV) C ondensed and sintered f luorapatite glass veneer 760 550 IPS e.max ZirPress (HV) H ot isostatically press (HIP) f luorapatite glass veneer 910 580 IPS e.max Press (LC) HIP l ithia disilicate glass ceramic core 930 560 IPS e.max ZirCAD (ZC) Yttria stabilized zirconia core 1500 1050* (15min) *Regenerating temperature Table 2 2. Dental ceramics and programmed firing parameters Ceramic Starting T(C) Heating rate ( C/min) Firing T (C) Holding t ime (min) Vacuum T on off ( C) GV 403 50 750 1 450 749 HV 700 60 910 15 500 910 LC 700 60 930 25 500 930 ZC* 1500 *fabricated by manufacturer

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30 Table 2 3. Elastic modulus, hardness, indent ation load, indentation strength (MPa) and fracture toughness [KIC] (MPam1/2) calculated from fractography (F) and indentation strength (IS) technique for dental ceramics Ceramics E (GPa) H (GPa) Indent load (N) Strength (MPa) K IC (F) (MPam 1/2 ) K IC (IS) (MPam 1/2 ) GV 65 5.4 9.8 48.7 3.0 0.76 0.04 0.83 0.04 HV 9.8 44.2 2.7 0.67 0.04 0.92 0.04 LC 96 5.5 19.6 154.5 20.9 2.66 0.42 2.45 0.25 ZC 208 13.1 196 193.8 1.8 4.31 0.11 5.12 0.03 Fig ure 21. Illustration of the determination of the glass transition temperature (Tg) of the LC group ( IPS e.max Press ) on the dilatometric curve. The rate of cooling is equal to 5 K/min. The in flection point corresponds t o a temperature of approximately 500 C.

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31 Fig ure 22. Schematic of a brittle material illustrating the critical flaw size [26] Figure 2 3. Optical microscopy ima ges of fractur e surfaces of bar specimens (a) yttria stabilized zirconia core ceramic (ZC) ; (b) lithia disilicate glass ceramic core (LC) ; (c) hot pressed glass veneer (HV) and (d) slurry condensed glass veneer (GV). Arrows demonstrates the critical crack boundaries.

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32 Figure 2 4. X ray diffraction pattern of an e.max Ceram veneering glass specimen (GV) Figure 2 5. X ray diffraction pattern of an e.max ZirPress veneering glass specimen (HV) Figure 2 6. X ray diffraction pattern of an e.max Press core ceramic specimen (LC)

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33 Figure 2 7. X ray diffraction pattern of an as sintered e.max ZirCAD core ceramic specimen (ZC) Figure 2 8. X ray diffraction pattern of a ZC core ceramic specimen after polishing An arrow indicate s the monoclinic peak of zirconia.

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34 Figure 2 9. X ray diffraction pattern of a ZC specimen after grit blasting by 100 m alumina for 1 min Figure 2 10. X ray diffraction patterns : ( a) sandblasted by 100 m alumina ; ( b) sandblasted after heat treatment at 1050 C for 15 min (t = tetragonal phase, m = monoclinic phase)

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35 Figure 2 11. SEM micrographs of e.max Ceram (GV) etched with an aqueous solution of 30% H2SO4 and 4% HF for 10 s : (a) 500x, and (b) 30000x Figure 2 12. SEM micrographs of e.max ZirPress (HV) etched with an aqueous solution of 30% H2SO4 and 4% HF for 10 s : (a) 5000x; (b) 20000x

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36 Figure 2 13. SEM micrograph (5000x) of e.max Press (LC) etched with an aqueous solution of 30% H2SO4 and 4% HF for 10 s Figure 2 14. SEM micrograph ( 30000x ) of e.max ZirCAD (ZC) thermally etched at 1420 C for 15 min

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37 CHAPTER 3 INTERFACIAL TOUGHNESS OF BILAYER DENTAL CERAMICS BASED ON A SHORT BAR, CHEVRONNOTCH TEST All ceramic restorations are desired by patients and dentists because of their excellent esthe tics and biocompatibility. Recently, new dental ceramic core systems with high strength and high toughness were introduced as alternatives to metals. However, chipping and fracture of veneer layers have been reported as the primary cause of failure for som e all ceramic systems [1422] which raises uncertainty on the benefits of using these all ceramic pros theses. A shear test is frequently used to fracture bonded ceramic materials and to measure the bond strength and stress at failure between bilayer structures Typically, analyses of s hear bond test data ignore the stress distribution generated within the adherence zone, which can have a significant effect on the magnitude of interfacial fracture stress and the mode of failure. Moreover, the loading device (blade or flat wire) for the shear test typically makes contact at a slight distance from the interfac e because of a small radius of curvature at the edge of the loading blade or width of the wire [62] This type of loading can cause flexural tensile failure rather than the desired shear failure. Flexural failure can lead to incorrect interpretation of the resultant bond strength data. Chevronnotch specimens have been developed for testing the fracture toughness of ceramics, high strength metals, and other brittle materials [6368] Barker proposed short rod and short bar designs for fracture toughness tests [63, 64, 66] These specimens have advantages over traditional fracture toughness specimens. They are relatively small, simple, and inexpensive. Precracks are not required because extremely high stress concentrations are produced upon loadi ng at the tip of the chevron notch. Thus, cracks can be initiated at a low applied load in a stable manner. Fracture toughness can be evaluated from the maximum load.

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38 Consequently, a loaddisplacement graph is not required and measurement of crack length i s not necessary [69] The primary advantage of the short bar fracture toughness test is that a precrack is not needed. Other advantages include the small sample dimension, the dependence of fracture toughness on the peak load alone, freedom from measuring the crack length, and the applicability to both ductile and brittle materials. Shortrod and short bar designs have been used to test the toughness of bonded interfaces between luting cements and dentin [70] and between resin composite and dentin [29] However, there are no published studies that have been based on the use of the chevronnotc h design f or testing of bilayer dental ceramics. Mecholsky and Barker [28] developed a method of using the chevronnotch specimens to measure the interfacial toughness of ce ramic metal interfaces. To ensure equal compliances for the two halves of the specimens with different elastic moduli, they assumed that, E1I1 = E2I2 where E1I1 and E2I2 are the product of the elastic modulus and moment of inertia of materials 1 and 2, res pectively. Values of d1 and d2 are chosen such that: E1d1 3 = E2d2 3 ( 31) where d is the thickness of the short bar. The compliances of both halves will be approximately the same (Fig ure 31a). The interfacial toughness can be derived from the equation; Kc = Ym P max / BW1/2 ( 32) where, Kc is the fracture toughness Ym is the minimum stress intensity coefficient Pmax is the maximum load to failure B is the width of the specimen W is the length of the specimen

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39 The obj ective of the present study was to determine the interfacial fracture toughness of two types of bilayer dental ceramics using the short bar chevron notch test and to test the null hypothesis that the interfacial toughness of core veneer bonded bilayers do not differ significantly from the fracture toughness of the monolithic veneer ceramic. 3.1 Materials and Methods 3.1.1 Material C ombinations Group LC/GV: Bilayer short bar specimens were prepared by bonding a lithia disilicate glass ceramic core (LC) to a fluorapatite glass veneer (GV) n=16. Group ZC/GV: Bilayer short bar specimens were prepared by bonding an yttria stabilized zirconia core (ZC) to a fluorapatite glass veneer (GV) n=16. Group GV/GV: Bilayer short bar specimens were prepared by bonding fluorapatite glass veneer to itself, n=16. 3.1.2 Sample P reparations Modified T shaped short bar core specimens (Figures 31b and 3 1c) of a lithia disilicate based core ceramic (LC) and an yttriastabilized zirconia core ceramic (ZC) [6.4 mm (B) x 11.5 mm (W) x 3 mm (d1)] were prepared using the lost wax technique and CAD CAM technology, respectively, according to the manufacturers i nstructions (Ivoclar Vivadent, Schaan, Liechtenstein) (Table 31). Palladium foil with a thickness of 25 m was placed on the core ceramic bonding surfaces and a V shaped notch was cut with dimensions as shown in F igure 31c to leave the bonding area expos ed (Figures 32a and 32b). The bonding surfaces were prepared according to the manufacturers instructions for each material. Veneering ceramic powder (GV) was incorporated with mixing liquid to obtain a slurry solution that was incrementally condensed on core ceramic substrates The thickness of veneering layers was adjusted according to equation 31. The modified T shaped short bar specimens (Figure 32c) were stored in 37 C distilled water for 30 days before testing to simulate oral aging conditions.

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40 For the control group, glass powder and buildup liquid were mixed and condensed in a silicone mold and sintered in a programmable furnace. The V shaped notch and another layer of glass veneer were prepared in the same manner. The chevron notches were prep ared using palladium foil. Instead of cutting the notch into the interface, the 25 m thick palladium foil was used before applying the glass veneer to create a thin, sharp notch. Because the toughness and strength of the core materials were much greater t han those of the glass veneer, cut notches were formed through the veneer layers, not between both layers across the interface. Chevronnotch specimens were tested using an Instron Universal Testing Ma chine (Model 5500R) at a cross head speed of 0.03 mm/mi n to promote stable crack growth until final failure occurred (Figure 3 2d). The fracture surfaces were individually examined using a magnifying lens and an optical microscope to determine if the crack initiated from the tip of chevron notch. In the presen t study, the elastic moduli were EZC = 165 GPa, ELC = 96 GPa and EGV = 65 GPa for the ZC, LC, and GV groups, respectively (Unpublished data, Ivoclar, Vivadent AG, Schaan, Liechtenstein). Specimens whose cracks propagated away from the interface were discarded and additional specimens were made to yield 16 specimens in each group. Two specimens from the control group were discarded because cracks propagated through the bulk material instead of along the interface. Figure 33 shows a fractured specimen from the LC/GV group. To determine the fracture toughness of the all glass veneer control group, eight glass veneer bars (2 mm x 4.5 mm x 28 mm) were prepared by mixing glass powder and buildup liquid, condensing in a silicone mold, and sintering according to the manufacturers instructions (Table 2 2) Indentation cracks were produced at the middle of each bar using a Vickers indenter at a 9.8 N load. The strength of each bar was determined 24 h after indentation using a four point flexure test with a 20.0 mm lower span and 6.7 mm upper span at a crosshead speed of 0.5

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41 mm/min. Fracture surfaces were examined and the fracture toughness of the glass veneer was calculated from the equation; KI c f c1/2 (3 3) where Y is a numerical constant that accounts for loading and crack geometry (1.65 for indentation cracks), c is the radius of an equivalent semicircular crack [equal to (ab)1/2] of depth a and half f is the stress at failure [26] The fracture toughness of the glass veneer was used to determine the dimensionless constant (Ym) for the chevron notch, short bar specimen with the geometry used in this study. By applying the mean load to failure (Pmax) and specimen dimensions (B, W) of fractured all glass veneer chevron notch specimens and the fracture toughness of the glass veneer control group (Kc) the dimensionless constant (Ym) was determined from equation 32. Statistical Analysis: The difference between the group mean s was analyzed using a one0.05 (SAS for Windows 9.1.3 service pack 4, SAS Institute Inc., Cary, NC, USA.). 3.2 Results The fracture toughness of the glass venee r, 0.74 (0.02) MPam1/2 was calculated using the fracture strength and fractography T he fracture toughness value of the glass veneer was used to determi ne the dimensionless constant, Ym The mean load s to failure of the LC/GV g roup and of the GV/GV c ontrol group were 38.9 (9.6) N and 41.9 (9.3) N, respectively, and were not significantly different (p>0.05). However, the mean load to failure of the ZC/GV Group was 7.1 (4.4) N, and this was significantly lower than the mean failure load of the other groups (p Table 32). T he interfacial toughness of the coreveneer was calculated from equation 32. The mean apparent interfacial toughness values of the LC/GV group and the ZC/GV group were approximately 0 .69 (0.11) and 0.13 (0.07) MPam1/2, respectively.

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42 3.3 Discussion Fracture toughness represents the resistance of a material to crack propagation. This is an important property for the design of dental prostheses that are usually made from brittle material s. Because dental ceramic restorations are usually made of bilayer or trilayer structures fracture toughness of the interface is also an important parameter in designing dental crowns and bridges, given that ceramic restorations often fail by chipping or delamination of the veneer layers [71, 72] Instead of calibrating the specimen geometry through the measurement of crack length on the chevron notch from the compliance calibration curve [29] or through finite element analysis [73] we obtained the value of Ym by performing measurement s on a material of known fracture toughness [63] In this case, we used the glass veneer bonded to the same glass veneer with a frac ture toughness derived from fracture strength and fractography as a control group to calculate the dimensionless constant (Ym), which is approximately 12 for thi s specimen config uration. Furthermore, the crosshead speed of 0.03 mm/min was selected to promote the required stable crack growth because a fast loading rate could cause cracks to propagate unstably. Figure 34 shows a loaddisplacement curve, which is ty pical of those from chevronnotch specimens in this study. The curve segment at the peak indicates stable crack growth before final failure (Figure 3 4) To create a crack along the interface, the thickness of each half was adjusted according to equation 31 to ensure equal compliance. Shown in Figures 35a and 35b are the SEM and optical microscopy images of the fracture surface features of LC/GV group specimens. The failures appear to travel along the interface. This study suggests that the mean interf acial toughness of the LC/GV group (0.69 MPam1/2) is comparable to the apparent fracture toughness of the GV/GV control group (0.74

PAGE 43

43 MPam1/2), indicating sufficient bonding. However, the mean interfacial toughness of the ZC/GV group (0.13 MPam1/2) was si gnificantly less than that of the other groups. This result suggests the need to critically evaluate the justification of using the same glass veneer on different types of ceramic cores under a simulated clinical situation. We found that the glass veneer t hat was hot pressed on the zirconia core yielded a greater mean interfacial toughness than the slurry condensed gl ass veneer on the zirconia core (Chapter 5). 3.4 Conclusions For bilayer allceramic restorations with high strength core materials, the veneering ceramics are the weakest link in the design of the structure. All ceramic restorations often fail from chipping of veneer layers or crack initiation at the interface [14, 15, 22] The high str ess concentrations within short bar chevronnotch specimens initiated a reliable stable crack. Fracture surface examination using optical and scanning electron microscopy indicated that cracks generally propagated wi thin the bonded interface. We conclude that the short bar, chevronnotch test is a useful method to determine the apparent interfacial fracture toughness of bilayer dental ceramics. Using a material with known fracture toughness as a reference material, th e dimensionless constant (Ym) could be determined without the need for compliance calibration or finite element analysis; however, the control group specimens are necessary and the constant obtained can only be applied to a specific geometry. Given that specimens have equivalent dimensions, the dimensionless constant could be determined and this technique provided a relatively reliable method for calculating the interfacial fracture toughness of the bilayer dental ceramics. The calculated apparent interfaci al toughness of the LC/GV group was 0.69 0.11 MPam1/2 and was comparable to the apparent toughness of the control group (GV/GV). Therefore, the null hypothesis was accepted. However, the apparent toughness of the ZC/G V

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44 group was 0.13 0.07 MPam1/2, wh ich was less than the apparent toughness of the control group. Thus the null hypothesis was rejected for this bilayer ceramic group. Table 3 1. Materials, ceramic types and processing t emperatures Ceramic Product Ceramic Type Processing Temperature IPS e.max Ceram (GV) Fluorapatite g lass veneer Sintered at 750 C IPS e.max Press (LC) Lithia disilicate based core ceramic Hot pressed at 930 C IPS e.max ZirCAD (ZC) YTZP zirconia core ceramic (4 6 wt% Y 2 O 3 ) Sintered at 1500 C Manufacturer: Ivoclar Vivadent, Schaan, Liechtenstein Table 3 2. Mean failure load and the apparent toughness; groups with the same superscript are not significantly different Ceramic groups Number of specimens Failure load (N) Interfacial Toug h ness (MPam 1/2 ) LC/GV 16 38.9 (9.6) a 0.69 (0.11) a ZC/GV 16 7.1 (4.4) b 0.13 (0.07) b GV/GV 16 41.9 (9.6) a 0.74 (0.17) a Fig ure 31. Specimen geometry: (a) A short bar geometry was modified for two dissimilar materials; (b) Schematic of crack plane of a specimen; and (c) Modified T shaped short bar specimen.

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45 Figure 32. Specimen preparation: (a) Core ceramic before veneering; (b) V shaped notch preparation; (c) Completed test specimen; and (d) Specimen in the test fixture. Fig ure 3 3. A fractured specimen of the LC/G V group

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46 Fig ure 3 4. Loaddisplacement curve of a chevron notch, short bar specimen Figure 3 5. (a) SEM image of fracture surface of chevron notch specimen of the LC/GV group, and (b) optical microscopy image of a fracture surface of a chevron notch specimen of LC/GV group

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47 CHAPTER 4 INTERFACIAL TOUGHNESS USING A CONTROLLEDFLAW MICROTENSILE TEST Shear and tensile bond tests have been used to test the bond strength between two bonded substrates. The commonly used shear bond test often produces fracture at a distance from the interfaces. This shift of the fractures at a distance from the interface is caused by misaligned loading and a nonuniform stress distribution [74, 75] and the geometry of the test specimen [76] and it may lead to an incorrect interpretation of the bond. Such noninterfacial fractures prevent the me asurement of interfacial bond quality because of the tendency toward cohesive fracture of one of the component materials and not by adhesive failure along the interface. Several research studies have been performed to measure the adhesion of resin bonding agents to ceramic [77, 78] usi ng a microtensile bond test. The microtensile bond test was originally designed to evaluate the bond strength between adhesive materials and dental tissues. This method minimizes the influence of interfacial defects because of the small bonding area of 1 m m2 or less [23, 79] Nevertheless, most results from microtensile bonding tests do not identify the location of failure origin so it is doubtful that the failure origins are necessarily located within the interfacial zone. Moreover, the strength data from microtensile tests vary with diffe rent specimen preparation methods and different sizes of specimens [25, 80, 81] Alumina based and zirconia based dental ceramics have attracted dentists to these products as the materials of choice as an alternative to metalceramic systems. However, the tougher and stronger core ceramics ar e veneered by weaker glassy phase ceramics to create color matched, translucent prostheses. Therefore, the weakest link among structural components is still the veneer layer or the interfacial region. Fracture and chipping of the veneer layers have been reported as the primary failures in all ceramic systems [17, 18, 82, 83]

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48 Thompson et al. [72] applied fracture surface analysis to examine clinically failed all ceramic (tetrasilicic fluoromica glass ceramic (TFGC) and magnesia aluminosilicate ceramic (MASC) crowns and compared these results with data from controlled laboratory test specimens fabricated from the same materials. They observed that TFGC crowns failed along the inner surface, and, for 78% of the MASC crowns, cracks initiated between the core ceramic and porcelain veneer interface. Failure stre ss can be estimated from the critical flaw size using fracture surface analysis combined with fracture mechanics principles. Kelly et al. [71] applied fractographic analysis to identify the failure origin of 20 laboratory failed and nine clinically failed all ceramic fixed partial dentures (FPDs). They found that the failure origin in 70 % of in vitro and 78 % of in vivo all ceramic FPDs occurred along the core veneer interface. Therefore, interfacial properties appear to play a major role in the performance of all ceramic dental restorations. The objective of the present study was to investigate the null hypothesis that a controlledflaw microtensile test of bilayer specimens can yield mean interfacial fracture toughness values that are not significantly different compared with those of monolithic ceramic control specimens. 4.1 Materials and Methods 4.1.1 Material Combinations Group 1: Bilayer bars of lithia disilicate glass ceramic core with a glass veneer (LC/GV) Group 2: Bilayer bars of yttria stabilized zirconia core with a glass veneer (ZC/GV) Group 3: Monolithic glass veneer bar (control group, GV/GV) 4.1.2 Specimen Preparation Lithia disilicate based core ceramics (LC) and yttriastabilized zirconia core ceramics (ZC) were prepared as 10 mm x 10 mm x 5 mm blocks. LC ceramic blocks were prepared by the lost wax technique according to the manufacturers instructions, and ZC blocks were prepared using CADCAM technology. The bonding surfaces were prepared as follows:

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49 Group 1: The surface of core ceramics (LC) was grit blast ed with Al2O3 at a pressure of 0.1 MPa and cleaned thoroughly with a hot stream of water before veneering pro cess Group 2: The surface of core ceramics (ZC) was only cleaned thoroughly with a hot stream of water before veneering process. Group 3: The surface of glass veneers (GV) was grit blast ed with Al2O3 at a pressure of 0.1 MPa, and cleaned thoroughly with a hot stream of water before veneering process The glass veneer was applied incrementally to the bonding area and fired in a programmable furnace (Programat P80, Ivoclar, Schaan, Liechtenstein). After firing, the grit abrasive on a metallographi c polishing disk, and the blocks were serially cut with a low speed diamond saw to obtain specimens with a 1 mm x 1 mm cross section for group LC/GV and a 1.5 mm x 1.5 mm cross section for group ZC/GV because the latter group tend to fracture during the cutting of specimens with a 1 mm2 cross sec tion (Figure 41). Specimens from the external surface were discarded to avoid the effect of surface stresses induced by polishing and thermal contraction. For the control group, glass powder and buildup liquid were mixed and condensed in a silicone mold and sintered in a programmable furnace. Glass veneer bars were cut in the same manner to obtain a 1 mm2 cross section for monolithic bar specimens. Before testing, cracks were initiated along the inter face of bilayer specimens using a Vickers indenter (Micromet 3, Buehler Ltd., IL, USA) at a load of 9.8 N at selected locations that were observed with an optical microscop e at a 4 00x magnification (Figures 42 and 4 3). Five adjoining indentations were m ade in each specimen to promote a crack along the interface. For the control group, cracks were produced in the center of the microtensile bars. Microtensile bars were attached to the flat grips using cyanoacrylate adhesive (Zapit, DVA, California, USA) an d tested in an Instron Universal Testing Machine (Instron 5500R) in a tensile mode at a cross head speed of 0.5 mm/min until fracture occurred (Figure 4 4). The specimens whose cracks did not initiate at the interface were rejected. This result was verified by

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50 using a fluorescent black light to discriminate the difference between core and veneer ceramics and an optical microscope to examine the fracture surface that initiated from the indentations along the interfaces. Load at failure was recorded, the stres s calculated, and fracture surfaces were observed for detection of critical flaws. Interfacial toughness (Kc) was calculated from equation 33. Fracture surfaces of microtensile bars were analyzed using optical microscopy and critical crack size images wer e captured by a digital camera attached to the optical microscope (uEye, model UI 1450C Obersulm, Germany ). Critical crack sizes were measured using computer software (Buehler Omnimet 8.81 Rev, Omnimet Modular Imaging System Buehler, Lake Bluff, IL ). A c alibration glass slide was used to verify the measurement at 100x, 200x and 400x magnifications. In addition, scanning electron microscopy was used to verify the size and location of the critical cracks at a g reater depth of field. Figures 4 5 and 46 show the optical micrograph and scanning electron micrograph respectively, of the critical flaw and location on a fractured specimen at 100x magnification. This technique was used to confirm that the failure initiated from the indent cracks that were placed al ong the interface. The arrows show the critical crack boundary and the original indents. 4.2 Results The mean, standard deviation (SD), and statistical analysis subsets of the apparent interfacial toughness values of the materials used in this study are lis ted in Fig ure 4 7 and Table 42. The interfacial toughness of the core/veneer interface of e.max Press/Ceram (LC/GV) and the apparent toughness of the monolithic control group ( GV/GV) were 0.63 0.13 MPa m1/2 and 0.70 0.13 MPa m1/2, respectively, and were not significantly different (p > 0.05; one way ANOVA and Bonferroni t tests). However, the difference in mean interfacial toughness of

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51 ZC/GV (0.37 0.12 MPam1/2) was significantly different from the mean values of the other two g roups (p 4.3 Discussion The mean tensile bond strength for bilayer materials is usually reported as the load at failure divided by the bonded ar ea, which is inversely proportional to the bonded surface area [25, 79, 81] Moreover, when the bond strength is greater than or equal to the tensile strength of the weakest substrate, the failure origins are likely to occur in either layer as a result of the great er probability of critical flaws in these layers. Furthermore, it is meaningless to report the bond strength values when the fracture origins are not identified sinc e the strength of the substrate is not known. A microtensile test was recently used to test the bond strength of zirconia core and veneering ceramics [84, 85] The failure types were mostly cohesive in nature. However, th e microtensile bond strength was not controlled by indent flaws and all types of failures (adhesive, cohesive, mixed) were included, which may not correspond to bond strength since cohesive failures represent the strength of substrates and are not representative of the interface. Furthermore, the microt ensile bond test alone might not be useful and should not be used when the strength of the bonding interfaces is relatively high because it causes cohesive failure of the substrate rathe r than of the bonded interface [86] In this study, we created indentation flaws along the interface to control the failure mode of the bonding ceramics. These flaws may be larger or sharper than the natural flaws in the s ubstrate. Therefore, interfacial failures can be produced and the interfacial toughness can be determined. This study applies a useful technique based on fracture mechanics principles and the microtensile test to calculate the apparent toughness of the cer amic interface. We expect that this

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52 technique will be independent of specimen dimensions based on previous fracture mechanics tests measuring the same toughness using different specimen dimensions. However, further research is ongoing to demonstrate that interfacial toughness of a bonded interface is a material property that is independent of the specimen cross sectional area. 4.4 Conclusions The interfacial toughness of the ZC/GV group was less than that of the LC/GV group and the apparent toughness of the monolithic glass veneer control group. This result suggests a need to evaluate the technique of bonding the veneer to the substrate. The controlledflaw microtensile test is a useful technique for calculating the interfacial toughness of bilayer dental ce ramics because the failure originates from a controlled flaw and propagates through the interface that represents the true bonding quality. Table 41. Materials, types, and processing t emperature Materials (Abbreviation) Type Process/ temperature IPS e.max Ceram (GV) Glass veneer Sintering/750 C IPS e.max Press (LC) Lithia disilicate based core ceramic Hot isostatic pressing/ 930 C IPS e.max ZirCAD (ZC) Y TZP zirconia core ceramic (4 6 wt% Y 2 O 3 ) Si ntering/ 1500 C Table 42. Mean interfacial toughness ( SD) and statistical subsets Ceramic Groups Number of specimens Toughness (MPam 1/2 ) Bonferroni subsets Z C/GV 17 0.37 0.12 B L C/GV 17 0.63 0.13 A GV/GV 16 0.70 0.13 A

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53 Figure 41. Sp ecimen p reparation Figure 4 2. Interfacial area of an LC/GV specimen

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54 Fig ure 43. An indent at the interface between core ceramic and veneer ceramic Fig ure 4 4. Specimen attached to the loading grips

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55 Figure 45. Optical micrograph of a critical flaw Arrows show the critical flaw boundary and original indents. Fig ure 46. SEM micrograph of same critical flaw as F ig ure 4 5

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56 Figure 47. Box plot of interfacial toughness determined from a controlled flaw microtensile test

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57 CHAPTER 5 EFFECT OF CROSS SECTIONAL AREA ON INTERFACIAL BOND STRENGTH AND INTERFACIAL TOUGHNESS OF BILAYER DENTAL CERAMICS All ceramic dental restorations have increased in popularity recently among dentists and patients because of the demand for esthetics and the biocompatibility concerns of metal exposure. H igh strength core ceramics, such as alumina and zirconia, were introduced as alternatives to metals. However, these high strength core ceramics are used as substructures for more translucent, but weaker glass or porcelain veneers. In addition, since all ceramic restorations are usually fabricated as bilayer structures, chipping of the ceramic veneer has been reported [1422] as a consequence of using these materials. Shear and microtensile bond tests have been used to test the bond strength of two bonded materials. However, the shear bond test often creates an initial cohesive crack that propagates away from the i nterface because of tensile stress caused by bending, i.e., a nonuniform stress dis tribution [74] and nonuniform circular cross sections [76] The microtensile bond test was first developed to dete rmine the bond strength between tooth structure and adhesive materials [87] and it has also been used to test the bond strength of core/ veneer ceramic systems [84, 85, 8890] The microtensile tes t produces a relatively uniform stress distribution. However, most of the previous tests did not use standard fractography methods to identify the failure origins. In addition, all failures have been generally included in the analysis of bond strength even though some of these may have failed cohesively. When the interfacial strength is compared to the cohesive strength of the strength of substrates, the failures tend to appear in the weaker substrate because of the greater probability of critical flaws in the larger volume of the substrate than along the interface [91] In addition, the microtensile bond strength may be inversely related to the cross sectional area of the test specimens [25, 7981]

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58 Because the bond strength can vary with specimen dimensions and flaw distributions, reported bond strengths can be misleading if smallsized specimens we re used compared with larger specimens. In contrast, fracture toughness should be independent of specimen dimensions. The microtensile bond test has been used to examine the bond strength between core and veneer ceramics [84, 88, 89] Although the mode of failure is sometimes classified by a percentage of cohesive or interfacial failure areas, the bond strength data are based typically on combining data from each of the two types of fracture. Fracture mechanics equations for the microtensile test design has been used to determine the apparent inte rfacial toughness of adhesive resin to ceramics [78] By using a controlled flaw along the interface, a crack is induced along the interface and the interfacial toughness can be calculated from the critical flaw size and the fracture strength. However, few studies have used this technique to determine the interfacial toughness of bonded bilayer ceramic str uctures. The objective of this study i s to test the hypothesis that the cross sectional area of a controlledflaw microtensile test has no effect on the interfacial toughness of bilayer dental ceramics in contrast to its effect on bond strength. 5.1 Materi als and Methods 5.1.1 Specimen Preparation Yttria stabilized zirconia core ceramic specimens ( IPS e.max ZirCAD, Ivoclar Vivadent AG, Schaan, Lichtenstein) were prepared from 10 mm X 10 mm X 5 mm blocks and were veneered using two different techniques, a hot isostatic pressing technique ( HIP, IPS e.max ZirPress ) and a condensed slurry technique ( CS, IPS e.max Ceram ) following procedures recommended by the manufacturer. The coefficients of thermal expansion between 100 C and 400 C for the zirconia ceram ic core, the press on glass veneer ( ZC/ HIP group), and condensed slurry glass veneer ( ZC/ CS group) are 10.8 ppm/K, 9.8 ppm/K and 9.3 ppm/K, respectively

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59 (Ivoclar Vivadent AG, Schaan, Liechtenstein) Bilayer ceramic blocks were finished by grinding with a 4 5 m grit abrasive on a metallographic polishing disk, and the blocks were serially cut using a low speed diamond saw. The sectioned bars had cross sectional areas of 1.0 mm2 and 1.4 mm2 for the ZC/ HIP group, and 1.4 mm2 and 2.3 mm2 for the ZC/CS group. 5.1.2 Interfacial Toughness Determination For the interfacial toughness test, a controlled flaw was created by a Vickers indenter along the interface using a 9.8 N load under ambient condition s Five adjacent indentation cracks were produced along the inte rface. The specimens were tested dry in tension at a crosshead speed of 0.5 mm/min. The interfacial toughness (Kc) was determined from the critical flaw size of the fracture surface and fract ure strength using equation 33. For the bond strength analysis, specimens were tested in tension without introducing a crack at the interface. In this study, only those specimens whose failures initiated at the interface were used in the interfacial toughness calculations 5.1.3 Microstructural Characterization of Inte rfacial Zones ZC/CS and ZC/HIP bilayer specimens (10 mm x 10 mm x 1 mm ) were embedded in epoxy resin as 10 mm cylinders. In addition, the same size of specimen was prepared for the lithiadisilicate glass ceramic bonded to condensed slurry glass veneer (LC /CS) to investigate the relationship of interfacial toughness to elemental diffusion between ceramic cores and glass veneers The embedded specimens were finished by grinding starting with 45 m grit metallographic polishing disk through 0.05 m alumina abrasive while exposed to a continuous flow of tap water. The specimens were sputter coated with conductive carbon in preparation for SEM/EDS analyses The composition of core and veneer ceramics was investigated using EDS (Energy Dispersive (X ray) Spectroscopy using a JOEL 6400 scanning electron microscope ( Jeol Ltd.,

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60 Tokyo, Japan) The detected elements from the EDS were used in interface analysis for possible migration of those elements through the interface from ceramic cores and glass veneers. The interface composition was analyzed by electron probe microanalysis using a line scan mode across the interfacial zone (EPMA JEOL Superprobe 733, Joel Ltd., Tokyo, Japan) and a ZAF correction method. Scanning electron probe analysis was performed from the glass veneer side, a cross the interface, and to the ceramic core. In addition, the interfacial zone of the lithia disilicate glass ceramic core bonded with condensed slurry glass veneer (LC/CS) was also investigated for comparison. The following parameters were u sed: acceleration voltage 15 kV; beam current 20 nA ; spot size 1 m ; scanned length 20 m ; step size 2 m ; and scanning time 10 s, for each element at each point. 5.2 Results Figure 5 1 shows original indents and a critical flaw of a fracture surface from a microtensile bar for the interfacial toughness test of the ZC/CS group. The mean interfacial toughness and the mean bond strength are shown in Table 51. A t test revealed that the cross sectional area has no effect on the interfacial toughness in both ceramic groups (p=0.8 and 0.9 for the ZC/HP and ZC/CS groups, respectively). In contras t, the mean bond strength varied with cross sectional area. In the CS group, the mean bond strength of specimens with a cross sectional area of 1.44 mm2 [26.0 ( 3.9) MPa ] was significantly greater than that for specimens with cross sectional areas of 2.25 mm2 [ 19.8 ( 2.8) MPa ] [p<0.0004]. However, the bond strength of the ZC/HIP gr oup could not be tested by this method since the specimens had a tendency to fail in the glass veneer layer when the interfacial toughness was relatively high compared with the fracture toughness of the glass veneer. In addition, the mean interfacial toughness of the ZC/HIP group was significantly great er than that of the ZC/CS group independent of cross sectional area [p<0.0001, one way ANOVA, Tukey's Studentized Range (HSD) Test].

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61 Figures 5 2 to 54 show the X ray spectra from EDS. The de tected elements were selected for elemental analysis by use of the characteristic X ray detection via electro n probe microanalysis (EPMA). The elements include d Zr, Si, Ca, P, Na, K, Hf, Al and Y. From the result s of EPMA, the interfacial zones were approximately 4 5 m wide in the ZC /HIP and ZC/CS groups (Figure s 55 and 56). No specific element s diffused beyond this zone in either group. Figure 5 7 shows the interfacial zone of the LC/CS group and interdiffusion zone of approximately 6 8 m was observed. 5.3 Discussion Shown in Figure 5 8 is an SEM micrograph that reveals a microtensile bar from the ZC/ HIP group with fewer and smaller pores than those from the ZC/ CS group The se pores may adversely affect the interfacial integrity, and they may have le d to the low interfacial toughness of the ZC/ CS group. The fracture strength of brittle materials such as ceramics is not generally reproducible. As flaws are already present in a material, there is a greater probability of finding a larger or more critica l flaw in larger size specimens than in those from smaller size specimens. In addition, when certain test conditions change, e.g., stressing rate, temperature, and testing environment, strength values also change [91] The strength of a given material depends on the size of the initiating crack present during processing in contrast to the fracture toughness, which is generally in dependent of the size of the initiating crack [33] Thus, it is more appropriate to employ fracture toughness as a mechanical property rather than the strength. In this study, we applied fracture mechanics principles to calculate the toughnes s of the interface. By applying a controlled flaw along the interface, interfacial failures were produced and the interfacial toughness was determined. In a pilot study, we observed that when the apparent toughness of interfaces was comparable to the fract ure toughness of the glass veneer, interfacial failure was less likely to occur in the bond strength test. Three cross sectional areas, 1 mm2 (1 mm x 1 mm),

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62 1.44 mm2 (1.2 mm x 1.2 mm), and 2.25 mm2 (1.5 mm x 1.5 mm), were selected for this study. However, the 1 mm2 specimens did not survive during the cutting process for the ZC/ CS group, and it was difficult to promote interfacial failure for the strength test in the 2.25 mm2 specimens for the ZC/ HIP group because of the ir highe r interfacial toughness. Sinc e the interfacial toughness of the ZC/ HIP group was comparable to the fracture toughness of the glass veneer, interfacial failure rarely occurred unless the controlled flaws were introduced at the interface. Thus, we did not perform the strength test as a f unction of cross sectional area on this group. The interfacial toughness of the ZC/ HIP group was greater than that of the ZC/CS group. For the ZC/ HIP group, the glass veneer was isostatically pressed on zirconia core ceramic at 91 0 C, and this glass vene er is more homogenous compared with the condensed slurry glass veneer. Some porous areas we re observed in HIP glass veneer. H owever those pores were distributed more uniform ly, and pores that were larger th an 50 m were rarely found (Figure 59). In c ontrast, pore size s in the CS glass veneer as large as 300 m were detected. In addition, the interfacial toughness of the ZC/HIP group was comparable to those of lithia disi licate core ceramic bonded to condense d slurry glass veneer (LC/CS) described in t he previous chapter. Electron Probe X ray Microanalysis (EPMA) is a quantitative elemental analysis method based on the generation of characteristic X ray s by a focused beam of energetic electrons [92] EPMA wa s used to analyze the concentration of element s (beryllium to actinides) at levels as low as 100 ppm. In this study, it was used in a line scan mode to analyze the elemental con centration along the interfacial zone. Interfacial zones 45 m in width were observed in both ZC/CS and ZC/HIP groups. The large r diffusion zone was expected for the materials with greater interfacial toughness ( ZC/H IP ). However, there was no difference i n diffusion depth between the ceramic cores and glass veneers for both bilayer composites. Therefore, the diffusion of the

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63 elements between zirconia core and glass veneers does not appear to control interfacial toughness in these two ceramic systems The d iffusion zone of approximately 6 8 m was observed in the LC/CS group. The interfacial toughness of the LC/CS was greater than that of the ZC/CS group as indicated in the previous study. Interestingly, the interfacial zone was rich in zinc compared with the amount in both the ceramic core and glass veneer Z inc was found in the wash coat that was applied on the lithia disilicate core before th e veneering process, but it was not found in Z irliner that was app lied on the zirconia core before veneering. In addition, diffusion of lithium which was the main composition in the lithia disilicate glass ceramic core could not be analyzed because of the limitation of EPMA to detect such a low a t omic number element. I n addition, fluorine which was stated to be in both types of glass veneer co uld not be identified using EDS. H owever, it was detected in small concentration using EPMA. In all ceramic systems, high failure rates give rise to the question of the overall benefit of these all ceramic materials. These failures may be caused by the thermal incompatibility stresses between ceramic cores and glass veneers that may introduce microcracks which can result in structural failure. It is generally thought that core m aterials should have a slightly greater [93] Under this condition, compressive axial residual stresses develop in outer veneer layers. However, if the positive mismatch is too large, not only do tangential compressive stresses develop in glass veneer, but radial tensile stresses may increase to great enough levels to also cause structural fracture [94] The effect of such a mismatch between ceramic cores and veneers was reported previously [95] In that study of a metalceramic system, it was found that a mismatch as great as or greater than 1 ppm/K could be deleterious to bond strength. In our study,

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64 the coefficient of thermal expansion between 100 C and 400 C of the zirconia core, press on glass veneer ( ZC/ HIP group) and slurry condensed glass veneer ( ZC/ CS group) were 10.8 ppm/K, 9.8 ppm/K and 9.3 ppm/K, respectively (Scientific Documentation, Ivoclar Vivadent Schaan, Liechtenstein ). The large mismatch between core and veneer (1.5 ppm/K) of the ZC/CS group may have caused the low interfacial toughness of this group. Nevertheless, the residual stresses in bilayer allceramic systems should be further investigated. In contrast to the bond strength, the interfacial toughness does not change with cross sectional area for the ceramic systems tested in this study. Thus, t he interfacial toughness is more reliable than bond strength for determining the bond quality between core and veneer ceramics. We as sume that this concept can be applied to other varieties of brittle materials as well. Table 5 1. Mean interfacial toughness and mean bond strength (SD) of ceramic groups with varying cross sectional area. Data sets for groups with the same superscript ar e not significantly different. Ceramic Group Cross section al area (mm 2 ) Interfacial toughness (MPa m 1/2 ) Bond strength (MPa) ZC/ HIP 1.00 0.64 [0.15] a (n=11) N/A 1.44 0.66 [0.08] a (n=9) N/A ZC/ CS 1.44 0.39 [0.09] b (n=13) 26.0 [3.9] A (n=11) 2.25 0.39 [0.11] b (n=14) 19.8 [2.8] B (n=11)

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65 Fig ure 51. SEM micrograph of a fracture surface of a microtensile bar in the ZC/CS group. White arrows indicate a critical crack boundary and black arrows point to original indentation sites Figure 5 2. Energy dispersive (X ray) spectrum of e.max Ceram veneering ceramic (CS )

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66 Figure 5 3. Energy dispersive (X ray) spectrum of e.max ZirCAD core ceramic (ZC ) Figure 5 4. Energy dispersive (X ray) spectrum of ZirPress veneering ceramic (HIP )

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67 Fi gure 55. X ray intensity of elements across the interface of the ZC/CS specimens from EPMA using the line scan mode

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68 Figure 5 6. X ray intensity of elements across the interface of the ZC/HIP specimens from EPMA using the line scan mode

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69 Figure 5 7. X ra y intensity of elements across the interface of the L C/CS specimens from EPMA using the line scan mode

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70 Figure 58. SEM micrographs of a microtensile bar from (a) the CS group (2.25 mm2), and (b) the HIP group (1.44 mm2) showing distribution of porosity F igure 5 9. P olished surface from (a) a HIP glass veneer specimen, and (b) a CS glass veneer specimen showing differences in the size and distribution of pores

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71 CHAPTER 6 RESIDUAL STRESS IN GLASS: INDENTATION CRACK AND FRACTOGRAPHY APPROACHES Residual stresses are often present in ceramic materials either by design or as a consequence of processing [96] Any manufacturing process that changes the shape of a glass phase ceramic or results in very large temperature gradients causes residual stress es to develop [97] In some cases, compres sive residual stresses are intentionally introduced to improve the mechanical property performance by inhibiting the propagation of crack s [98101] Regardless of the source of residual stress, it plays an important role in serviceability of a component. A variety of techniques have bee n used to determine the magnitude of residual stresses, and each method has some limitations. For example, X ray or neutron diffraction techniques can be applied effectively only to crystalline materials. Birefringent methods or the photoelastic technique are limited to transparent materials, and electricalstrain gauge applications are limited in their ability to distinguish residual stresses from the total stress [97, 102] The crack indentation technique based on a pyramidal diamond indenter has been widely used to measure mechanical properties of brittle materials such as hardness, brittleness, and f racture toughness [36] This technique has the potential for rapid evaluation of material properties of small samples and may be useful for materials development or quality control [103] Recently, the Vickers indentation method was criticized as an unreliable test for fracture toughness determination or for any other fracture resistance parameter [32, 104] Additionally, calibration constants have been empirically determined and adjusted according to the material under investigation to achieve reasonable results [32] However, because of its expediency, convenience, and small sample requirement, this technique still retain s its popularity as a qualitative tool to determine mechanical parameters of ceramic materials.

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72 Indentation methods have been proposed since the 1930s as means of measuring or detecting surface residual stresses [97] The effective compressive or tensile residual stresses can be estimated by comparing the mean fracture stress of stressed sample s to that in unstressed samples The indentation technique has also been employed to calculate global residual stresses in bilayer dental ceramics [105] These global stresses are calculated using a fracture mechanics equation combined with local residual stresses fro m contact damage. A new indentation method was proposed to measure the residual stresses adjacent to a Vickers indentation crack [106] and this method can be extended to other cases. T o determine the residual stress adjacent to an i ndentation, an indentation must first be made in unstressed material, and the stress intensity factor for the indentation is calculat ed from the equation [107] : KC E H 1 2 P c0 3 2 ( 61) where Kc is the critical fracture toughness, P is the indentation load, c0 is the crack length at that influence of residual stress, the crack grows to a new equilibrium length (c) at the same indentation load (P) as is used in th e stress free case. At equilibrium, the crack will experience a composite stress intensity, and Kc is given by the equation: KC E H 1 2 P c3 2 ac1 2 ( 62) Combining equations R) can be calculated using equation 63: R KC1 c0c 3 2 c1 2 ( 63)

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73 median crack, the surface crack is assumed to be semicircular in shape, and has a value of 1.26. The crack length in the unstressed material is c0, and c is the crack length under the influence of the residual stress. A similar method that involves the superposition of stress intensity factors was used to measure near surface residual stresses in tempered glass based on another fracture mechanics approach [108] and was also used to calculate the local residual stress in bilayer dental ceramics [105] This method employs a beam having a surface residual stress that is subjected to flexural loading. For a semi elliptical crack of depth a and half wi dth b, c is the equivalent semi circular crack size, c, where c = (ab)1/2. The stress intensity factor at th e border of the semicircular crack is calculated from the following equation: KI 2 1 2 Ac1 2 YF 2 1 2 Rc1 2 YR ( 64) where YF rection factor associated with flexural loading, and YR the crack border correction factor associat ed with the residual stress field [108] In the case of f lexural loading, YFR unity. Residual stress can be calculated as follows: R2 1 2 Ac1 2 YF KIYR 2 1 2 c1 2 ( 65) We will test the hypothesis that there is no significant differen ce between the mean residual stresses determined from the indentation crack technique and those determined from fracture surface analysis

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74 6.1 Materials and Methods 6.1.1 Specimen P reparation Glass bars (2.3 mm x 4 mm x 28 mm) were prepared from a soda lime silicate glass plate. All sharp edges were beveled by grinding with a 30 remove high stress concentration areas. The glass transition temperature (Tg) of soda lime silicate g la ss was determined using DSC (differential scanning calorimetry) at a heating rate of 3 C/min. All specimens wer e annealed at 650 C (75 C above Tg) for 30 min to eliminate all residual stresses before testing. 6.1.2 Fracture Toughness T est Twelve beam specimens were indented using a Vickers hardness indenter at a load of 19.6 N while being viewed under an optical micros cope at 400x magnification (Figure 6 1). Crack measurements were made within 1 min and 24 h after indentation using a calibr ated imaging system (OmniMet Modular Imaging System, Buehler, Lake Bluff, IL) at 100x magnification. The mean crack lengths measured within 1 min and 24 h after the indentation were chosen as the crack size of the unstressed materials (c0) to calculate the residual stress from equation 63. Indent s pecimens were divided into two groups. For Group 1, the annealed and indented specimens were subjected to four point flexure to failure with a 20.0 mm lower span and 6.7 mm upper span at a crosshead speed of 0.5 mm/min. For Group 2, the specimens were annealed at 650 C for 30 min after indentation to remove the residual stress from the indent prior to flexure testing Failure strength was calculated, the critical crack size was measured, and the fracture toughnes s (Kc) was determined from the equation: Kc = Y f c1/2 ( 66) where Y is a numerical constant that accounts for location, loading condition and crack geometry [1.65 for indentation cracks (Group 1) and 1.26 for indentation cracks with local residual s tresses

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75 from the indentation being removed (Group 2)], c is the equivalent semi circular critical flaw size calculated from (ab)1/2, where a is the semiminor axis, and b is the semimajor axis of a semielliptical crack [26] Crack indentation approach : A total of 18 specimens were heated at 650 C for 30 min and compressive residual stress was intentionally introduced within the surface by applying compressed air at a pressure of 0.05 MPa a t a controlled distance of 2.5 cm from the center of the 18 bars for 1 min after removing the specimens from the furnace (Figure 6 1) Indentation cracks were produced at the middle of each bar under a 19.6 N load using a Vickers hardness indenter. Crack m easurements were made on each bar within 1 min (Group A) and 24 h (Group B) after indentation. The crack length, which is perpendicular to the long axis of the bar, was measured while this crack was subjected to tensile loading (Fig ure 62 ). The mean crack length (c) from this group was compared wi th the mean crack length (c0) of the annealed group within 1 min and 24 h after indentation. Residual stresses were calculated using equation 63. Fractography approach : A total of 18 specim ens from above were tes ted using a fourpoint flexure test (Group C). Fracture strength was calculated, fracture surfaces were analyzed, and residual stresses of each bar were calculated using equation 65. 6.1.3 Statistical A nalysis Residual stresses were calculated using equa tion 63 were based on crack size measurements made within 1 min (Group A) and 24 h (Group B) after indentation and from equation 65 for each specimen in Group C. The stresses were analyz ed for significant differences using a randomized block design. Beca use all measurements were made on the same specimens and were compared among each other using different approaches, the variability within each specimen (within each block) was less than that between specimens (between

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76 blocks). Statistical analyses were pe rformed at a significance level of 0.05 (SAS 9.1.3 Service pack 4). 6.2 Results Mean f racture toughness and stren gth of the soda lime glass are listed in Table 61. Fracture surfaces were analyzed to determine critical crack dimensions (Figure 63). Although the differences in mean strength between Group 1 and Group 2 were significantly different, the difference in fracture toughness was not statistically significant regardless of the method used in this study [p=0.62]. The fracture toug hness in Ta ble 61 was used in the determination of surface residual stresses in the soda lime silica glass. Indentation cracks of unstressed and stre ssed specimens are shown in F igure 64. The average crack size in the surface of glass bars without residual stress w as larger than that of bars with residual compressive stress for the same indentation load. The calculated residual s tresses are shown in Table 6 2. Compressive residual stresses, which were calculated using the crack indentation technique for Groups A and B, were not significantly different from those determined using the quantitative fractographic approach (Group C) [p=0.19 and 0.13, respectively]. However, the difference in mean residual stress between Group A and Group B was significantly different [p=0 .003]. We could not measure the crack length of four specimens from Group A because no cracks initiated during that period of time. However, the crack lengths of the same specimens were observed at 24 h (Group B). 6.3 Discussion Table 62 shows the residual stress es of Group A and Group B that were calculated using equation 63 compared well with those calculated using equation 65 (Group C). There was no significant difference between both techniques; the results show good agreement between the indentatio n technique at 24 h (Group B) and the fractographic approach (Group C). Because of

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77 the greater effect of slow crack growth during the early period after indentation, the crack measurements of Group A were more difficult to obtain than those of Group B as r apid crack propagation occurred during the early period. Thus, a computer controlled, rapid imaging system was connected to the indenter and optical microscope. Furthermore, cracks were not well developed in some specimens of Group A, and they could not be observed (four specimens), but they could be detected at 24 h. The techniques presented in this study were based on fracture mechanics principles for determining the compressive residual stress i n glass surfaces. Both techniques are similar in that they involve the superposition of stress intensity factors. For indentation cracks, the Vickers indenter was used with a moderate load and a permanent indentation was formed with cracks of approximately equal radial crack lengths from each corner of the impress ion. Residual stress was determined by comparing crack lengths of indentation with a residual stress field, with those crack lengths from indentations in the unstressed material using the same indentation load [109] C racks from the materials with a tensile stress field surrounding the crack are longer than those in the unstressed materials, whereas cracks under a compressive stress field will be shorter. For four point flexure with a surfaceindented crack, the applied loading results in a flexural stress A) in the surface of the bars (Fig. 6.5 ) due to the superposition of stresses, and the critical stress intensity at the c rack tip in this case will be greater than the critical stress intensity factor of the unstressed material [108] This method is useful for determining surface residual stresses. For the indentation technique, it is essential to understand the residual stress distribution within materials [31, 106, 109] However, the indentation technique requires a high quality polished and flat specimen s urface; otherwise, the crack length cannot be correctly measured.

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78 In conclusion, there w as no significant difference between residual stresses calculated using the indentation technique and the fractographic approach based on induced surface cracks, four point flexure, and fracture surface analysis. Therefore, the indentation technique may be employed under carefully controlled conditions as a simplified method to determine the effective residual stress value within brittle materials such as glasses, and glass ceramics. Table 61. Fracture strength (MPa) and fracture toughness (MPa m1/2) of so da lime glass used in this study Condition Strength (MPa) Fracture Toughness (MPa m 1/2 ) Anneal + Indent (Group 1) 47.6 2.7 0.82 0.07 Anneal + Indent + Anneal (Group 2) 64.7 2.1 0.81 0.04 Table 62. Residual stress (MPa) of each specimen from the indentation technique and the superposition of the str ess intensity factor method (* N o crack observed) Specimen # Residual stresses from heat treatment calculated from indentation technique (MPa) Residual stress from heat treatment calculated from fractography (MPa) Group A (1 min) Group B (24 h) Group C 1 6. 7 5.8 7.3 2 11.5 15.1 13.2 3 15. 3 22.7 23.4 4 N/A* 29.8 16.9 5 N/A 24.9 25.6 6 N/A 17.9 13.3 7 12.3 20.6 2.1 8 23.3 22.3 13.5 9 12.7 23.1 23.5 10 18.5 20.2 14.2 11 23.3 27.3 28.3 12 N/A 31.5 31.7 13 13.5 20.6 19.4 14 14. 4 21.9 24.0 15 14 0 15.8 16.5 16 14 0 20.2 19.3 17 7.0 10.3 12.4 18 8 0 12.3 11.7 Mean (SD) 13.9 ( 5.2) 20.1 ( 6.6) 17.6 ( 7.6)

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79 Figure 61. Schematic of the design of heat strengthening apparatus [Adapted from Anusavice et.al ., 1989 [110] ] Fig ure 62. Schematic of bar specimen showing an indentation crack used for residual stress calculation.

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80 Figure 63. Optical microscopy images of the fracture surface showing a critical crack of a fracture toughness test specimen (a) with local residual stress from an indent (Group 1) and (b) without local residual stress form an indent (group 2) [annealed after indentation]. Ar rows denote the critical crack boundary. Fig ure 64. Optical microscopy images of an indentation crack induced at a load of 19.6 N in (a) an annealed specimen and (b) a specimen with compressive residual stress. The crack length s of specimens wit h compressive residual stress are shorter.

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81 Figure 65. S chematic of a semicircular crack with indentation crack depth [a] within the global R] under uniform flexural A] [108]

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82 CHAPTER 7 RESIDUAL STRESSES IN BILAYER ALL CERAMIC SYSTEMS H igh strength and toughness core ceramic materials such as zirconia and alumina, were introduced as alternative materials to metals to support weaker, but more translucent glass veneer. All ceramic restorations are gain ing more popularity among patients and dentists because of their superior biocompatibility, chemical stability and esthetics. However, chipping or delamination of the glass veneer layers have been reported as causes of failure [1 422] In addition, the results from two clinical studies that used fractography to determine the location of fracture origin indicat ed that the core veneer interface was an important site of failure [71, 72] Those failures may be controlled by the most critical flaw present in the restorations or residual stresses generated from thermal incompatibility. Residual stresses are usually present in the structure of bilayer dental ceramic systems because of the difference in thermal expansion/contraction between core and veneer ceramics [2] In metal ceramic systems, it is generally suggested that the thermal expansion/contraction of glass veneer and metal should be closely matched T he expansion/contractio n of metal should be slightly greater than the glass veneer so that the glass veneer is under compressive stress on cooling from the firing cycle [111] This is based on the fact that glass veneer is much weaker in tension, and the state of compression would enhance the loadbearing capability of the restorations [112] However, not only are compres sive hoop and axial stresses generated, but the tensile radial stresses also develop in the veneer layers when thermal contraction of the core is greater than that of the glass veneer [93] If high residual tensile stresses develop in all ceramic systems, th ey may fail under service in the oral cavity [46] Moreover, this may cause a high incidence of cohesive failure within ceramic veneers or lessen the bond strength between core and veneer ceramics [75, 113]

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83 In the previous chapter, the indentation technique and fractography were used to determine the residual stresses with i n the surface of soda lime silica glass. In this study, the same approaches were used to investigate the residual stresses i n bilayer dental ceramic systems and determine the possible relationship between the in terfacial toughness and residual stresses developed in these systems. The specific aim of this study wa s to test the hypothesis that the indentation technique yields no difference in sur face residual stresses compared with the fractographic approach. 7.1 Materials and Methods 7.1.1 Specimen P reparation 7.1.1.1 Monolithic bars control groups Six bar specimens (4.5 mm x 2 mm x 28 mm) wer e prepared for each group. For the IPS e.max Ceram glass veneer group [GV], powders were mixed as slurry and condensed in a polyoxymethylene mold and fired according to manufacturers recommendations For the IPS e.max ZirPress hot isostatic ally pressed glass veneer group (HV), selfcured acrylic resin (Pattern resin GC Corp., Tokyo, Japan) was used to prepare the HV bar spe cimens. The resin bars were sprued and invested with IPS PressVest special investment material (Ivoclar Vivadent AG, Schaan, Liechtenstei n). The invested bars were burned out for 60 min at 850 C in a preheating furnace (Radiance, Jelrus Int., Hicksville, NV, USA) and the mold cavity was filled by hot pressing IPS e.max ZirPress ingots using a pressing furnace (EP 500, Ivoclar Vivadent AG, Schaan, Liechtenstein) at 910 C. The cooled specimens were removed from the mold and divested with 80 m glass beads ( Williams glass bead, Ivoclar Vivadent North America, Amherst, NY) at a pressure of 0.2 MPa. All bar specimens were polished from 45 m alumina abrasive through 0.05 m abrasive (Mark V Laboratory, East Granby, CT, U.S.A.). All the edges were rounded to minimized stress concentration. The GV group and HV group were annealed at

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84 550 C and 580 C respectively to remove all the residual stresses. These groups are used as the cont rol groups for residual stress calculation s 7.1.1.2 Bilayer bar specimens Six c ore ceramic specimens (1 mm x 4.5 mm x 28 mm) wer e prepared for each group. For lithia disicate glass core ceramics (LC), six specimens were prepared from wax patterns (Rio Grande NM, USA). Wax patterns were sprued, invested as mentioned above, and hot pressed with IPS e.max Press ingot at 930 C. For yttria stabilized zirconia core ceramics (ZC), presintered IPS e.max ZirCAD ceramic block was cut using a low speed diamond saw (Buehler Isomet, Lake Bluff, IL, USA) to obtain 12 bar specimens and sintere d at 1500 C by the manufacturer. The bars of the LC group and six bars of the ZC group were veneered with IPS e.max Ceram glass veneer and sintered according to the manufacturers instructions (Table 2 2) The remaining bars of the ZC group were veneered with the press on glass veneer (IPS e.max ZirPress ). Each specimen was adjusted and polished from 45 m alumina abrasive through 0.05 m abrasive while exposed to a continuous flow of tap water. The following veneer and core/veneer specimens were prepared: Group 1: GV Group 2: HV Group 3: LC/GV Group 4: ZC/GV Group 5: ZC/HV The fina l dimension of the specimens were 2 mm (thickness) x 4.5 mm (width) x 25 mm (length). The core/veneer thickness ratio was 1:1. 7.1.2 Testing M ethods Fracture mechanics was used to calculated residual stresses in the veneer layers of bilayer dental ceramics (Group 3 5). The following steps were used:

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85 1. Indentations were made on bilayer specimens and monolithic specimens and crack sizes of both groups were compared. 2. Failure load was measured and flexural strength of each group was calculated. 3. Apparent fracture toughness of the bilayer bars were determined using fracture mechanics equation s 4. Residual stresses in bilayer specimens were determined using flexural strength and fractography. I ndentation crack s were induced i n the glass veneer surfaces using a Vickers hardness indenter at a load of 9.8 N Crack measurement was made at 200 X using a calibrated imaging system (OmniMet Modular Imaging System, Buehler, Lake Bluff, IL, USA) at 24 h after indentation to allow for complete growth of the cracks caused by the effect of the environment and residual contact stresses. Cracks which were perpendicular to the long axis of the specimens were measured because these axes corresponded to the cr acks under tensile loading (Figure 7 1). Crack sizes of bilayer bars were compared with those of the control groups (Group 1 & 2) of the same glass veneer s for calculating the residual stresses using the indentation technique from the equation: R KC1 c0c 3 2 c1 2 (7 1) median crack, the surface crack is assumed to be semicircular in shape, and has a value of 1.26. The crack length in the unstressed material is c0, which is the ave rage crack size of annealed monolithic specimens and c is the crack length of bilayer bar under the influence of the residual stress from thermal incompatibility f) of each group was determined using four point flexure test with a 20.0 mm support span and a 6.7 mm loading span at a crosshead speed of 0.5 mm/min. The veneer surface was placed in tension for all specimens. The flexure strength of the bilayer bar specimens were calculated using composite beam theory (Appendix A) [114]

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86 The f racture surface of each specimen was observed using an optical microscope at 100 X magnification. Selected specimens were examined under SEM. Only specimens whose failures originated from the indentation were used in the study. Critical crack size was measured to determine the fracture toughness of each specimen. Fracture toughness (Kc) was derived usi ng: Kc = Y f c1/2 (7 2) where Y is a geometric factor that accounts for location, loading condition and crack geometry ( 1.65 for indentation cracks ) and c is the equivalent semi circular critical flaw size calculated from (ab)1/2, where a is the semiminor axis, and b is the semimajor axis of a semielliptical crack [26] Resid ual stresses calculated from flexural strength and fractography were calculated from the following equation : R2 1 2 Ac1 2 YF KIYR 2 1 2 c1 2 (7 3) where YFR the crack border correction factor as sociated with the residual stress field [108] In the ca se of flexural loading, YFR unity. 7.1.3 Statistical Analysis Oneway analysis of variance and multiple comparisons using the Tukey Kramer method were performed to determine whet her there are differences among group means for flexural strength crack size, apparent fracture toughness and residual stresses of specimens in each group. Additionally, residual stress calculated from the indentation technique (equation 71) and from th ose of flexural stresses and fractography (equation 7 3) were compared using complete

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87 randomized block design. A statistical analysis was performed at a significance level of 0.05 using SAS 9.1.3 Service pack 4. 7.2 Results Figure 7 2 shows a fracture of a bilayer bar specimen of the LC/GV group. The critical crack was found to originate from an indentation introduced on the tensile or glass veneer surface of the bilayer specimen. The mean apparent fracture toughness (Kc), fl exural strength f) crack length, apparent fracture toughness, and the residual stresses R) calculated from the indentation crack and fracture stress and fractography are presented in Table 7 1. One way analysis of variance shows that there were statistically sign ificant diffe rences among the means of flexural strength crack length and apparent fracture toughness (p 0.0007). For the flexural stre ngth, there was no significant difference between f of the GV and the HV groups. Additionally the fract ure stress f, of the monolithic glass veneer did not differ from that of the bilayer specimens using identical glass veneer s (p=1.0) except for the ZC/GV group whose f was significantly less than that of the GV group (p = 0.006). The mean crack length o f the ZC/GV group was greater than those of other groups (p 0.007) However, there were no significant difference between the mean crack length s of the bilayer specimens on the zirconia core (p = 1.0) No significant difference in Kc was found between two monolithic glass veneers used in this study (p = 0.64). In addition, Kc of monolithic specimens did not differ from those of bilayer specimens using identical glass veneers (p > 0.05). However, Kc of the LC/GV was greater than that o f the other bilayer specimens (p A fracture surface of a bilayer bar of the ZC/GV group i s shown in F igure 73. The SEM images were captured in specimens whose crit ical crack boundaries were questionable because of the depth of fie ld of the SEM im ages. Figure 7 4 shows

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88 optical and SEM images of a fracture surface of a specimen from the LC/GV group. The critical crack originated from an indentation created on the tensile surface of bilayer bar specimen. R values are summarized in Tab le 7 2. Residual stresses R) on the surface of glass veneers calculated from the crack indentation technique were not significantly different from those determined using the quantitative fractographic approach (p = 0.23). The mean R of the LC/GV group w as different from those of the ZC/GV group for both techniques (p < 0.05) and the R on the surface of glass veneers on the ZC cores were not different regardless of veneering technique (p < 0.05). 7.3 Discussion The observation of the fracture surfa ces showed that all the failure origins were produced by the indentation cracks that were introduced as a controlled flaw on the veneer surface. A controlled flaw reduced the variation of flexural strength of the specimens [115] For the LC/GV group, cracks originated from flaws i n the surfaces of the veneer layer and propagated through the cores However, when a high toughness zirconia core ceram ic as was used to support the glass veneer, cracks propagated and arrested at the interfaces regardless of the glass veneer type D elamination occurred b efore the final failure of the cores at a greater stress levels However, in this study, the first failure was considered as the failure of the structures. Similar results were found for cracks that initiated from the glass veneer and tended to arrest and deflect at the interfaces as they propagated to ward high toughness core materials [116] In contrast to the result s of a previous study, this crack propagation behavior showed that the flexural strength of the bilayer specimens was greater than that of monolithic specimens [105] In this study, t he mean f of monolithic glass veneer specimens did not differ from that of the bilayer specimens using same glass veneers except for the ZC/GV group whose strength was less than that of the GV group. In addition, the mean crack length in the ZC/GV group was greater than that of the

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89 other groups which indicates that the residual tensile stress near the surface of the glass ven eer corresponds to a lower mean f than that of the glass veneer control group (GV). Moreover, the interfacial toughness may play an important role for the lower strength value. The apparent toughness ( Kc) of the bilayer specimens did not differ from those of the matching glass veneer control group, which indicates that the most critical flaw induced in the glass veneer was subjected to the greatest tensile stress, and it controlled the failure of the composite structures. However, a finite element analysis showed that the internal surface of the glass ceramic crown was subjected to the largest tensile stress under occlusal loading [117] therefore, high toughness core substrate materials may be beneficial Moreover, the role of core support on the failure of bilayer structures was investigated. The results of this analysis showed that a stiffer core, especially alumina, can be useful as an inhibitor of cone crack failure in the veneer by suppressing tensile stresses close to the interface [118] Nevertheless, such advantages could be compromised by the residual stresses from thermal incompatibilities between cores and veneers [118, 119] Residual stresses of the ZC/GV group were different from those of the LC/GV group regardless of the technique used. Although the same glass veneer was employed, the R in th e ZC/GV group were tensile in n ature compared with that of the L C/GV group while they were compressive The residual stresses in the ZC/GV group corresponded to decreased strength s longer crack length s and a lower Kc when the residual tensile stresses were developed in the veneer layer. In addition, these stresses may have severely affect ed the bond quality between ceramic cores and glass veneers [75, 113] It was found that the apparent interfacial toughness of the ZC/GV bilayer was l ess than that of the LC/GV group ( Chapter 4). This difference may be because of the greater thermal incompatibility between the ZC and GV ( 1.5 ppm/K ) than that of

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90 the LC/GV group (1 ppm/K) as reported by the manufacturer. However, these residual thermal in compatibility stresses should be further investigated. Although there were no differences between f, crack length, Kc, and R between the ZC/GV and ZC/HV groups, the former group had the lower interfacial toughness than that of the latter group. In this case, both the thermal incompatibility and venee ring technique may be important role in bonding integrity. In the present study, the monolithic glass veneer groups (GV and HV) were annealed and used as control groups. It was assumed that there was no residual stress after the annealing process. The calculated residual stress of the LC/GV group from the indentation technique (E quation 71) and fractography (Equation 73 ) ranged betwe en 11 MPa and 7 MPa, and between 21 MPa and 3 MPa respectively. For the ZC/GV group, R values were between 7 MPa and 10 MPa, and 2 MPa and 16 MPa respectively. For the ZC/HV group, they ranged from 1 MPa to 9 MPa and 6 MPa to 17 MPa respective ly. The range f or the LC/GV group was different from that in another study for materials of a similar composition whose residual stresses ranged from 40 MPa and 60 MPa. However, these compressive residual stresses had to be offset by the tensile strength and such high level of st resses seem to be high compared with the tensile strength of the glass veneers. 7.3 Conclusions Wh en c ritical cracks are introduced in glass v eneer layers and those layers a re subjected to a high tensile stress, the flexural strength and apparent fracture toughness are controlled primarily by the mechanical properties of the glass veneer layers. These properties apparently improve after compressive residual stresses a re generated from thermal incompatibilities. However, it sho uld be keep in mind that when compressive residual stresses are intentionally introduced in to the glass veneer, those compressive stresses are counterbalanced by tensile stresses within the glass veneers themselves at the interfaces, or in the core cerami cs.

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91 Table 7 f), crack length, apparent fracture toughness (Kc), and residual stresses ( R) calculated from indentation cracks and fractography of monolithic and bilayer specimens (#) = standard deviation Groups f [MPa] Crack length [m] K c [MPam 1/2 ] fractography R [MPa] indentation R [MPa] fractography GV 48.7 (3.0) a b 185 (9) A 0.76 (0.04) xy z N/A N/A HV 44.1 (2.5) bc 189 (9) A 0.66 (0.04) wy z N/A N/A LC/GV 49.8 (3.6) a 184 (19) A 0.85 (0.12) x 1.3 (6.8) Y 8.2 (9.0) C ZC/GV 40.2 (2.6) c 216 (7) B 0.68 (0.08) wz 8.2 (1.4) Z 6.0 (7.6) D ZC/HV 41.1 (5.6) c 204 (16) A B 0.61 (0.10) w 3.6 (3.4) YZ 5.7 (10.0) D Letter in t he superscripts illustrate the significan t difference between groups. Table 7 2. Residual stresses R (MPa) of each specimen based on the indentation technique and fractography Specimen # Group R (indentation) R (fractography) 1 LC/GV 6.6 3.8 2 LC/GV 5.9 0.3 3 LC/GV 2.6 6.7 4 LC/GV 4.3 3.2 5 LC/GV 3.5 16.7 6 LC/GV 7.1 21.4 7 LC/GV 11.3 12.4 8 ZC/GV 7.9 9.3 9 ZC/GV 7.9 13.3 10 ZC/GV 7.0 0.5 11 ZC/GV 6.6 0.4 12 ZC/GV 10.3 1.9 13 ZC/GV 9.3 15.6 14 ZC/HV 9.0 9.7 15 ZC/HV 0.9 6.4 16 ZC/HV 2.8 3.7 17 ZC/HV 3.7 16.7 18 ZC/HV 1.4 2.1 19 ZC/HV 5.3 16.1

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92 Figure 7 1. Schematic of specimen dimensions of bilayer specimens showing indentation crack used for calculation of residual stress. Figure 7 2. A fracture surface of a bilayer specimen of the LC/GV group. Arrow demonstrates the critical flaw from a n original indent.

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93 Figure 7 3. F ract ure surface of a specimen from the ZC/GV group showing a critical crack boundary (black arrows) and the original indent ( white arrow) Figure 7 4. C ritical flaw on the fracture surface of a bilayer specimen from the LC/GV group ( a) optical micrograph and ( b) SEM micrograph. Black arrows outline the critical crack boundary.

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94 CHAPTER 8 CONCLUSIONS All ceramic restorations, especially in posterior regions, are usually fabricated as bilayer structures. H igh strength and toughness core ceramics, such as alumina, zirconia or alumina zirconia composites which are opaque, have been used as substructures to support more esthetic and translucent, but weaker, glass veneer layers. Compared with metal ceramic sys tems, no longterm clinical studies have been reported for all ceramic restorations and most of the studies were within the past 10 years Because the failures of all ceramic restorations have been frequently reported as chipping or delamination of the gla ss veneer layers, interfacial properties seem to play important roles in such failures. For the interfacial toughness tests, this study used two different approaches, a s hort bar chevronnotch test and the controlledflaw microtensile test. The results of those techniques are summarized in Table 8 1. The present results are in contrast to those from conventional bonding tests such as shear bond tests or traditional microtensile tests that usually result in failures at a distance from the interfaces [24] In addition, the conventional bond strength test values are inversely proportional to their cross section areas. A chevron notch test and a co ntrolled flaw microtensile test can c reate int erfacial failures that represent true bonding quality and are independent of the cross sectional area. This consistency in results occurs because the most critical cracks were formed from sharp triangular shapes of the chevron notch or from adjoining indentations in the microtensile tests. The results show that the interfacial toughness of the ZC/GV group was less than that for the other groups. This may raise the concern of using such veneering techniques on high toughness zirconia. In addition, the resul ts from both techniques seemed to be in agreement. However, the value of the interfacial toughness using the short bar chevron notch specimens

PAGE 95

95 seems to be less than that of the controlledflaw microtensile test in the ZC/GV group. An explanation for this r esult is that because the chevron notch test specimens were kept in 37 C distilled water fo r 30 days before testing but the controlledflaw microtensile tests were tested in ambient conditions. It is evident that stabilized zirconia is susceptible to low temperature degradation such as by water or other aqueous solutions over the range of 65500 C and this effect is more catastrophic at lower temperatures and in shorter times in aqueous environments [47, 120] This degradation occurs because of the unfavorable transformation from the tetragonal to monoclinic phas e that occurs especially on the surface of zirconia, which is more unstable than the bulk. However, the potential of such degradation in the oral environment and the effect of temperature should be further studied. Residual stresses ( R) usually develop in ceramic restorations because of the thermal incompatibility between ceramic or metal cores and glass veneers. A variety of techniques can be used to determine the magnitude of R in the structures. The indentation technique has been used to determine the R in brittle materials by comparing the crack size differences between a stress free material and material with residual stresses. However, the indentation technique has been criticized for its lack of accuracy in determining the stress intensity facto r. In this study, we measured the R of soda lime silica glass and bilayer dental ceramic using the indentation technique compared with that using flexural strength and fractography, and found that there was no statistically significant difference between those techniques. The study showed that the indentation technique can be used as a simplified method for estimating the residual stresses in glass. However, there are limitations of this technique; specimens require a highquality polished and com pletely flat surface in order that precise indentations can be made and crack lengths can be accurately measured.

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96 Interfacial toughness and residual stress ( R) in bilayer dental ceramics may be inter related. In this study, when a glass veneer was used on different core ceramics, the group with the greater interfacial toughness (LC/GV) had residual compressive stresses in the glass veneer while residual tensile stresses were found in the glass veneer with weaker interfacial toughness (ZC/GV). Howeve r, there was no difference in R between the ZC/GV and ZC/HV groups, but greater in terfacial toughness was found for the ZC/HV group. In this case, pores in the glass veneer or the thermal incompatibility most likely play an important role in the bond quality The hot isostatic glass veneer (HV) on zirconia core was less porous than the condensed slurry glass veneer (GV). In addition, the difference s in the thermal expansion coefficients between the core ceramic and glass veneer of the ZC/GV and ZC/HV groups we re 1 .5 ppm/K and 1 ppm/K respectively. Such a high mismatch may reduce the stress needed to cause bond failure However, it is unlikely that the thermal incompatibility alone will be sufficient to cause interfacial failure [121] On the contrary, it is difficult to identify the one or two principal factors that a re reliable enough to predict the level of thermal incompatibility. Moreover, there was an additional thin layer of liner that was applied on the core before the veneering process a nd this further increased the complexity of this system The introduction of high toughness and strength core ceramics has created broad interests in dentistry. However, it is clear that long term clinical studies are necessary to determine whether they ar e practical systems for restorations especially in posterior regions that are subjected to high loads

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97 Table 8 1. Interfacial toughness of bilayer specimens ; groups with the same superscript are not significantly different Ceramic groups Interfacial t oughness (MPam 1/2 ) Chevron notch test Controlled flaw microtensile test GV/GV (control) 0.74 0.17 A 0.70 0.13 a LC/GV 0.69 0.11 A 0.63 0.13 a ZC/GV 0.13 0.07 B 0.37 0.12 b ZC/HV N/A 0.64 0.15 a

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98 APPENDIX A FLEXURAL STRENGTH CALCULATED FROM COMPOSITE BEAMS Beams that are fabricated from more than one material are called composite beams. The flexure formula for monolith ic beams cannot be applied to the beams that consist of more than one material. Therefore, the flexural stress on the tensile surface of bilayer dental ceramics was determined from an approximate theory used in the bending of composite beams [114] ,i.e., the transformed section method [114, 122] x) along the composite beam varies linearly from top to bottom and can be express by: x = (A 1) According to Hookes law, the normal stress distance y from the n eutral axis can be expressed as: x1 = E1 x2 = E2 (A 2) x1 x 2 are stress in material 1 and 2 respectively. E is elastic modulus. For the theory for bending of composite beams, because the neutral axis does not pass through the centroid of the cross sectional area, the neutral axis is found under the condition that the resultant axial force actin g on the cross section equals to zero; therefore, 1x 1dA 2x 2dA 0 (A 3) where the first and second integral is evaluated over the cross sectional area of material 1 and 2 respectively. Since the curvature is constant at any given cross section, the equation c an be converted to another form: E11y dA E22y dA 0 (A 4)

PAGE 99

99 The moment of curvature can be derived from; 1I1 + E2I2) (A 5) By substituting equation A 5 in equation A 2, the stress in material 1 and 2 can be obtained from: x 1 MyE1E1I1 E2I2 X 2 MyE2E1I1 E2I2 (A 6) For the transformationsection method, this approach transforms the cross section of a composite beam into an equivalent cross section of an imaginary beam of one material ( Fig. A 1). This new cross section is a transformed section. This transformed section depends on the ratio of elastic modulus of the adjoining beam (n), where n = E2/E1 and E1 and E2 are elastic modulus of material 1 and 2 respectively. Assuming that E2 > E1, the neutral axis of the new cross section is obtained from the equation A 3 which can be rewritten as; E11y dA E22y n dA 0 (A 7) Since the transformed beam consists of only one material, stress in the transformed beam can be given by standard flexure formula; x 1 M y IT (A 8) where IT is the moment of inertia of the transformed section with respect to neutral axis in which; IT = I1 + nI2 = I1 + (E2/E1) I2 (A 9) Substituting A 9 into A 8 gives the s ame result as A 6; x 1 MyE1E1I1 E2I2 Stresses in material 2 are not the same and must be multiplied by the elastic modulus ratio n to obtain the stresses in material 2.

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100 x 2 M y IT n Stresses in material 2 are the same as equation A 6. x 2 MyE2E1I1 E2I2 In this study, stress calculation using composite beam theory was demonstrated on bilayer dental ceramics. A bilayer bar of e.max Press core (LC) /e.max Ceram veneer (GV) with a 1 mm: 1 mm ratio was subjected to a 50N failure load and the GV layer sustained tensile stress The elastic modulus of LC and GV are 96 GPa and 65 GPa respectively; therefore, the modulus ratio of the LC/GV = 96/65 = 1.48. The f irst step is to locate the neutral axis of the cross section. Let h1 and h2 be the distances from the neutral axis. 1 y dA = (h1 0.5 mm) (4.5 mm x 1 mm) 2 y dA = (1.5 h1) (4.5 mm x 1 mm) = (h1 1.5 mm) (4.5 mm) According to equation A 4, E11y dA E22y dA 0 65 GPa (h1 0.5 mm) (4.5 mm) + 96 GPa (h1 1.5 mm) (4.5 mm) = 0 h1 = 1.0963 The m oment of inertia of areas 1 and 2 (I1 and I2) can be obtained using the parallel axis theorem [114] I1 = (1/12) (4.5 mm) (1 mm)3 + 4.5 mm (1 mm) (h1 0.5 mm)2 = 1.995 mm4 I2 = (1/12) (4.5 mm) (1 mm)3 + 4.5 mm (1 mm) (h2 0.5 mm)2 = 1.095 mm4 I = (1/3) (4.5 mm) h1 3 + (1/3) (4.5 mm) h2 3 = 3.09 mm4, or = I1 + I2 Strength from the bilayer bar that failed at a 50 N failure load wa s calculated from the following:

PAGE 101

101 f = MyE1/(E1I1+E2I2) = (PL/6)(1.1 mm)(65 GPa)/(65 GPa.995 mm4)+(96 GPa.095 mm4) = 50.75 MPa For the transformed section method, the first bilayer beam will be transformed into a beam of material 1 according to the modular ratio, and the new width of the second layer (where E2 > E1) = 4.5 x (E2/E1) to 4.5 x 1.48 mm or 6.65 mm u sing the bottom edge as a reference line, and with distance yi measured upward. Since the transformed section consists of only one material (Fig. A 2), the distance h1 can be calculated as follow: h1 = i Ai / i 0.5 m m 4.5 m m 1 m m 1.5 m m 6.65 m m 1 m m 4.5 m m 1 m m 6.65 m m 1 m m = 1.1 mm For the moment of inertia of the transformed section, the m oment of inertia of the entire cross section with respect to the neutral a xis can be calculated as follow: IT = (1/12)(4.5 mm)(1 mm)3 + (4.5 mm)(1 mm)(h10.5 mm)2 + (1/12) (6.65 mm)(1 mm)3 + 6.65 (h20.5)2 = 3.612 mm4 Under a 50 N load to failure, the strength of the monolithic bar under four point flexure with a 20.0 mm support span and 6.7 mm loading span was calculated f rom the following: f = My/I = PL/bh2 = (50 N)(20 mm)/(4.5 mm)(2 mm)2 = 55.6 MPa For a bilayer bar with 1 mm: 1 mm ratio, the failure strength was calculated from: f = My/IT = PLy/6 IT = (50 N)(20 mm)(1.1 mm)/6(3.612 mm4)= 50.76 MPa

PAGE 102

102 Fig ure A 1. Cross section schematic of the bilayer beam of the LC and GV. Figure A 2. Transformed section schematic of the bilayer beam of the LC and GV

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BIOGRAPHICAL SKETCH Dr. Chuchai Anunmana was born in Bangkok, the capital city of Thailand, in 1970. After high school, he passed the national dental school entrance examination and was admitted to Faculty of Dentistry, Mahidol University, Bangkok, Thailand in 1987 and received a Doctor of Dental Surgery (D.D.S.) degree in 1993. He subsequently received a Graduate Diploma of Clinical Sc ience in Prosthodontics from Mahidol University in 1996. From 1993 to 1995 he worked as a lecturer at the Department of Prosthodontics, Chiang Mai University, where he met his wife, Siriwan He worked in a private dental practice as a prosthodontist until 2002 and became a facult y member in the Department of Prosthodontics, Faculty of Dentistry, Mahidol University. He earned a scholarship from the Thai Government for graduate studies and was admitted to the graduate program in Material s Science and Engineering at the University of Florida. Dr. Chuchai Anunmana and his wife Siriwan, an endodontist, enjoy their life with their beautiful children, daughter Pam and son Pete. Upon receiving his Doctor of Philosophy degree, Dr. Chuchai Anunmana will resume his work as a faculty member and research scientist in the Department of Prosthodontics, Faculty of Dentistry, Mahidol University.