Applied Science, Faculty of
Engineering, School of (Okanagan)
DSpace
UBCO
Robles Arellano, Karen Denisse
Master of Applied Science - MASc
University of British Columbia
Development of new technological systems, such as fuel cells or nuclear power generation stations, often relies on the availability of materials whose properties accommodate the operational requirements of the system. Fabrication of creep-resistant ceramic materials is currently receiving worldwide interest due to their potential use in solid oxide fuel cells, thermal barrier coatings and insulators for hostile environments. In this research a novel fabrication process, Spark Plasma Sintering (SPS), was used to prepare cubic Yttria-stabilized Zirconia (YSZ) composites doped with three different rare earth oxides. Specifically, YSZ was doped with 6.4wt%, 12.7 wt% and 18.8 wt% CeO₂; 11.6 wt%, 21.6 wt% and 30.5wt% La₂O₃; and 12wt% and 22 wt% Nd₂O₃. Composites were successfully fabricated using the SPS technique, and the effect of selected parameters on density and grain size was studied. The novel composites were characterized using Scanning Electron Microscope –X-Ray Energy Dispersive Spectroscopy (SEM-XEDS), X-Ray Diffraction (XRD) and Optical Microscopy (OM). Also, compressive creep testing was performed and the active creep deformation mechanisms were studied. The results reveal that increasing the sintering temperature resulted in grain growth in all the composites. Addition of the rare earth oxides had a strong grain refinement effect on YSZ. Further, addition of La₂O₃ and Nd₂O₃ resulted in formation of a bimodal microstructure. In contrast, the effect of variation of dopant wt% on the grain size and relative density of the composites could not be established. Compressive creep tests in the temperature range of 1200-1330°C were carried out for YSZ, 6.4wt% CeO₂–YSZ and 11.6wt% La₂O₃ –YSZ composites. The stress exponent for all composites ranged between 1.2-2.7. The activation energy was ~730 kJ/mol for YSZ, ~490 kJ/mol for 6.4wt% CeO₂–YSZ and ~566 kJ/mol for 11.6wt% La₂O₃ –YSZ. These values of n and Q suggest that creep deformation was primarily accommodated by grain boundary sliding possibly assisted by lattice diffusion. These results were supported by microstructure evaluation, where evidence of grain boundary sliding was observed as crenellated cracks along boundaries, as well as grain boundary cavitation. There was a potential stress-induced phase transformation in the YSZ material, possibly related to the metastable t’ or monoclinic phases.
DEVELOPMENT, CHARACTERIZATION AND CREEP RESISTANCE OF CUBIC ZIRCONIA DOPED WITH RARE EARTH OXIDES PROCESSED THROUGH SPARK PLASMA SINTERING by Karen Denisse Robles Arellano B. A. Sc., The University of British Columbia, 2011 A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF MASTER OF APPLIED SCIENCE in THE COLLEGE OF GRADUATE STUDIES (Mechanical Engineering) THE UNIVERSITY OF BRITISH COLUMBIA (Okanagan) April 2013 © Karen Denisse Robles Arellano, 2013 ii Abstract Development of new technological systems, such as fuel cells or nuclear power generation stations, often relies on the availability of materials whose properties accommodate the operational requirements of the system. Fabrication of creep-resistant ceramic materials is currently receiving worldwide interest due to their potential use in solid oxide fuel cells, thermal barrier coatings and insulators for hostile environments. In this research a novel fabrication process, Spark Plasma Sintering (SPS), was used to prepare cubic Yttria-stabilized Zirconia (YSZ) composites doped with three different rare earth oxides. Specifically, YSZ was doped with 6.4wt%, 12.7 wt% and 18.8 wt% CeO2; 11.6 wt%, 21.6 wt% and 30.5wt% La2O3; and 12wt% and 22 wt% Nd2O3. Composites were successfully fabricated using the SPS technique, and the effect of selected parameters on density and grain size was studied. The novel composites were characterized using Scanning Electron Microscope –X-Ray Energy Dispersive Spectroscopy (SEM-XEDS), X-Ray Diffraction (XRD) and Optical Microscopy (OM). Also, compressive creep testing was performed and the active creep deformation mechanisms were studied. The results reveal that increasing the sintering temperature resulted in grain growth in all the composites. Addition of the rare earth oxides had a strong grain refinement effect on YSZ. Further, addition of La2O3 and Nd2O3 resulted in formation of a bimodal microstructure. In contrast, the effect of variation of dopant wt% on the grain size and relative density of the composites could not be established. Compressive creep tests in the temperature range of 1200-1330°C were carried out for YSZ, 6.4wt% CeO2–YSZ and 11.6wt% La2O3 –YSZ composites. The stress exponent for all composites ranged between 1.2-2.7. The activation energy was ~730 kJ/mol for YSZ, ~490 kJ/mol for 6.4wt% CeO2–YSZ and ~566 kJ/mol for 11.6wt% La2O3 –YSZ. These values of n and Q suggest that creep deformation was primarily accommodated by grain boundary sliding possibly assisted by lattice diffusion. These results were supported by microstructure evaluation, where evidence of grain boundary sliding was observed as crenellated cracks along boundaries, as well as grain boundary cavitation. There was a potential stress-induced phase transformation in the YSZ material, possibly related to the metastable t’ or monoclinic phases. iii Preface A version of Chapter 4 Section 2 has been submitted for publication. [K. D. Robles Arellano], L. Bichler, K. Akkiraju, R. Fong (2013) Fabrication and Properties of Spark Plasma Sintered CeO2-YSZ Ceramic Composite. Materials Characterization (under review). Sample preparation was conducted by K. Akkiraju and S. Ipe Varghese at Indian Institute of Technology, Madras, and M. van Hanegem at University of British Columbia. I conducted all of the imaging and result analysis, and wrote most of the manuscript. A version of Chapter 4 Section 3 has been submitted for publication. [K. D. Robles Arellano], L. Bichler, K. Akkiraju, R.Fong, K. Mondal (2013) Densification behavior of Spark Plasma Sintered La2O3-YSZ Ceramic Composites. Ceramics International (under review). Sample preparation was conducted by K. Akkiraju and M. van Hanegem at University of British Columbia. I conducted all of the imaging and result analysis. The manuscript was prepared in collaboration with L. Bichler and K. Mondal. iv Table of Contents Abstract .................................................................................................................................... ii Preface ..................................................................................................................................... iii Table of Contents ................................................................................................................... iv List of Tables ......................................................................................................................... vii List of Figures ......................................................................................................................... ix List of Abbreviations ........................................................................................................... xiv List of Symbols ...................................................................................................................... xv Acknowledgements .............................................................................................................. xvi Dedication ........................................................................................................................... xviii Chapter 1: Introduction ......................................................................................................... 1 Chapter 2: Literature Review ................................................................................................ 4 2.1 Introduction to Creep Resistance ............................................................................................ 4 2.2 Creep Mechanisms ................................................................................................................. 7 2.1 Zirconia Ceramics .................................................................................................................. 9 2.2 Addition of Cerium Oxide (CeO2) to Zirconia ..................................................................... 10 2.3 Addition of Lanthanum Oxide (La2O3) to Zirconia .............................................................. 13 2.4 Addition of Neodymium Oxide (Nd2O3) to Zirconia ........................................................... 15 2.5 Overview of Spark Plasma Sintering of YSZ ....................................................................... 16 2.6 Chapter Summary ................................................................................................................. 20 Chapter 3: Experimental Procedure ................................................................................... 21 3.1 Precursor Powders ............................................................................................................... 21 3.2 Precursor Powder Blending .................................................................................................. 22 3.3 Spark Plasma Sintering ......................................................................................................... 23 3.3.1 Sintering of YSZ and CeO2-YSZ Composites ................................................................ 23 3.3.2 Sintering of La2O3-YSZ and Nd2O3-YSZ Composites ................................................... 25 3.4 Polishing, Thermal Etching and Density Measurement ....................................................... 28 3.5 Microscopy and XRD ........................................................................................................... 29 v 3.6 Compressive Creep Experiments.......................................................................................... 29 3.6.1 Sample Preparation ......................................................................................................... 29 3.6.2 Creep Apparatus ............................................................................................................. 30 3.6.3 Creep Testing Procedure ................................................................................................. 31 3.6.3.1 Constant Crosshead Speed Test ................................................................................ 31 3.6.3.2 Varying Load and Temperature Test ........................................................................ 33 3.6.4 Derivation of n and Q ..................................................................................................... 34 3.6.4.1 Calculation of n ................................................................................................... 34 3.6.4.2 Calculation of Q .................................................................................................. 35 Chapter 4: Results and Discussion ...................................................................................... 38 4.1 YSZ ...................................................................................................................................... 38 4.1.1 Densification ................................................................................................................... 38 4.1.2 Chemical Analysis and Microstructure ........................................................................... 40 4.1.3 XRD Analysis ................................................................................................................. 40 4.1.4 Discussion ....................................................................................................................... 41 4.2 YSZ doped with CeO2 .......................................................................................................... 42 4.2.1 Densification ................................................................................................................... 42 4.2.2 Chemical Analysis and Microstructure ........................................................................... 44 4.2.2.1 6.4 wt% CeO2-YSZ composite sintered at 1300 °C ............................................. 44 4.2.2.2 12.7 wt% CeO2-YSZ composite sintered at 1300 °C ........................................... 46 4.2.2.3 18.8 wt% CeO2-YSZ composite sintered at 1300 °C ........................................... 48 4.2.2.4 18.8 wt% CeO2-YSZ composite sintered at 1400 °C ........................................... 50 4.2.3 XRD ................................................................................................................................ 52 4.2.4 Discussion ....................................................................................................................... 53 4.3 YSZ doped with La2O3 ......................................................................................................... 55 4.3.1 Densification ................................................................................................................... 55 4.3.2 Chemical Analysis and Microstructure ........................................................................... 57 4.3.2.1 11.6 wt% La2O3-YSZ composite sintered at 1400 °C .......................................... 57 4.3.2.2 11.6 wt% La2O3-YSZ composite sintered at 1500 °C .......................................... 59 4.3.2.3 21.6 wt% La2O3-YSZ composite sintered at 1500 °C .......................................... 61 4.3.2.4 30.5 wt% La2O3-YSZ composite sintered at 1500 °C .......................................... 65 4.3.3 XRD Analysis ................................................................................................................. 69 4.3.4 Discussion ....................................................................................................................... 70 4.4 YSZ doped with Nd2O3 ........................................................................................................ 72 vi 4.4.1 Densification ................................................................................................................... 72 4.4.2 Chemical Analysis and Microstructure ........................................................................... 73 4.4.2.1 12 wt% Nd2O3-YSZ composite sintered at 1400 °C............................................. 73 4.4.2.2 22.2 wt% Nd2O3-YSZ composite sintered at 1400 °C.......................................... 75 4.4.3 XRD Analysis ................................................................................................................. 77 4.4.4 Discussion ....................................................................................................................... 78 4.5 Compressive Creep Results .................................................................................................. 78 4.5.1 YSZ ................................................................................................................................. 79 4.5.1.1 Microstructure of YSZ after creep ........................................................................ 80 4.5.2 6.4 wt% CeO2-YSZ ........................................................................................................ 85 4.5.2.1 Microstructure of 6.4 wt% CeO2-YSZ after creep ............................................... 86 4.5.3 11.6 wt% La2O3-YSZ ..................................................................................................... 92 4.5.3.1 Microstructure of 11.6 wt% La2O3-YSZ after creep ............................................ 93 Chapter 5: Conclusions ........................................................................................................ 98 5.1 SPS Processing ..................................................................................................................... 98 5.2 Microstructure Characterization ........................................................................................... 98 5.3 Creep Performance ............................................................................................................... 99 Chapter 6: Suggested Future Work .................................................................................. 101 6.1 Optimization of SPS-processed Zirconia Composites ........................................................ 101 6.2 Characterization of Rare Earth-doped Zirconia Composites .............................................. 101 References ............................................................................................................................ 103 Appendices ........................................................................................................................... 110 Appendix A: Sample Calculations for n and Q ........................................................................ 110 A.1 Setup ................................................................................................................................. 110 A.2 Variables Obtained from Testing Parameters and Constants ............................................ 110 A.3 Variables Obtained From Direct Measurements ............................................................... 111 A.4 Variables Obtained From Calculation .............................................................................. 111 A.5 Calculating n and Q ......................................................................................................... 112 Appendix B: Micrographs ......................................................................................................... 120 vii List of Tables Table 1 Chemical composition of precursor powders…………………………………….. 21 Table 2 Levels of doping of YSZ with rare earth oxides…………………………………. 23 Table 3 SPS processing parameters for YSZ and CeO2-YSZ composites………………... 23 Table 4 SPS processing parameters for La2O3-YSZ and Nd2O3-YSZ composites……….. 26 Table 5 Polishing procedure for microstructure evaluation………………………………. 29 Table 6 Representative procedure for a VLT test………………………………………… 34 Table 7 Effect of SPS parameters on the grain size and relative density of YSZ Composite………………………………………………………………………... 38 Table 8 Chemical composition of matrix in YSZ composite…………………………….. 40 Table 9 Effect of SPS parameters and doping level on the grain size and relatie density of CeO2-YSZ composites………………………………………………... 43 Table 10 Chemical composition of features in 6.4 wt% CeO2-YSZ composite sintered at 1300 °C, 30 Mpa, 5 min………………………………………………. 45 Table 11 Chemical composition of features in 12.7 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min……………………………………………… 47 Table 12 Chemical composition of features in 18.8 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min……………………………………………… 49 Table 13 Chemical composition of features in 18.8 wt% CeO2 – YSZ composite sintered at 1400 °C, 30 MPa, 7 min……………………………………………… 52 Table 14 Effect of SPS parameters and doping level on the grain size and relative density of La2O3 – YSZ composite………………………………………………. 56 Table 15 Chemical composition of features in 11.6 wt% La2O3 – YSZ composite sintered at 1400 °C, 30 MPa, 3 min……………………………………………… 59 Table 16 Chemical composition of features in 11.6 wt% La2O3 – YSZ composite sintered at 1500 °C, 50 MPa, 10 min…………………………………………….. 61 Table 17 Chemical composition of features in 21.6 wt% La2O3 – YSZ composite sintered at 1500 °C, 50 MPa, 10 min…………………………………………….. 63 Table 18 Chemical composition of features in 30.5 wt% La2O3 – YSZ composite sintered at 1500 °C, 50 MPa, 10 min…………………………………………….. 67 viii Table 19 Effect of doping level on the grain size and relative density of Nd2O3 – YSZ composite ……………………………………………………….... 72 Table 20 Chemical composition of features in 12 wt% Nd2O3 – YSZ composite sintered at 1400 °C, 40 MPa, 5 min ……………………………………………... 75 Table 21 Chemical composition of features in 22 wt% Nd2O3 – YSZ composite sintered at 1400 °C, 40 MPa, 5 min …………………………………………....… 76 Table 22 Chemical composition of features in YSZ composite before and after creep…..... 82 Table 23 Chemical composition of features in 6.4 wt% CeO2-YSZ composite before and after creep ……………………………………………………...…….. 87 Table 24 Chemical composition of Cerium-rich particles in post-creep 6.4 wt% CeO2-YSZ composite …………………………………………………..………... 91 Table 25 Chemical composition of features in 11.6 wt% La2O3 – YSZ composite before and after creep ……………………………………………………………. 93 ix List of Figures Figure 1 Logartithmic plots used to estimate: a) n, b) Q, c) p………………………………. 6 Figure 2 Lattice diffusion mechanisms…………………………………………………….... 8 Figure 3 Grain boundary sliding schematic showing void formation……………………….. 8 Figure 4 Crystal structure of Zirconia……………………………………………………….. 9 Figure 5 Phase diagrams: a) Ceria-Zirconia binary system, b) Ceria-Yttria-Zirconia system at 1400 °C in Zirconia-rich region…………………………………...…… 11 Figure 6 Phase diagram for Lanthana-Zirconia system……………………………………. 13 Figure 7 Phase diagram for Neodymia-Zirconia system developed by Wang et al……...... 15 Figure 8 SPS apparatus schematic…………………………………………………………. 18 Figure 9 SEM micrographs of cubic Zirconia prepared with: a) SPS, b) conventional sintering…………………………………………………………. 19 Figure 10 SEM micrograph of Scandia-stabilized Zirconia prepared with SPS…………….. 19 Figure 11 SEM micrographs of precursor powders: a) YSZ, b) CeO2, c) La2O3, d) Nd2O3………………………………………………………………... 22 Figure 12 Representative samples of as-sintered: a) YSZ sintered at 1300 °C, 30 MPa, 7 min, b) YSZ sintered at 1400 °C, 30 MPa, 7 min, c) sintered at 1500 °C, 50 MPa, 10 min…………………………………………. 24 Figure 13 Representative samples of as-sintered: a) 6.4 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min, b) 12.7 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min, c) 18.8 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min, d) 18.8 wt% CeO2 – YSZ composite sintered at 1400 °C, 30 MPa, 7 min…………………………………………….… 25 Figure 14 Representative samples of as-sintered: a) 11.6 wt% La2O3 –YSZ composite sintered at 1400 °C, 30 MPa, 3 min, b) 11.6 wt% La2O3 –YSZ composite sintered at 1500 °C 50 MPa, 10 min, c) 21.6 wt% La2O3 –YSZ composite sintered at 1500 °C 50 MPa, 10 min, d) 30.5 wt% La2O3-YSZ composite sintered at 1500 °C 50 MPa, 10 min…………………………………………….... 27 x Figure 15 Representative samples of as-sintered: a) 12 wt% Nd2O3 –YSZ composite sintered at 1400 °C, 40 MPa, 5 min, b) 22.2 wt% Nd2O3 –YSZ sintered at 1400 °C, 40 MPa, 5 min……………………………………………………….….. 28 Figure 16 Samples for creep testing: a) Sample location within the as-sintered disc, b) dimensions of creep specimen……………………………………………….… 30 Figure 17 Compressive creep apparatus: a) front view of the machine, b) schematic of the compression chamber …………………………..…………………………….. 31 Figure 18 Stress vs. strain curve for YSZ at 1300ºC and a deformation rate of 1μm/min…... 32 Figure 19 Schematic of the VLT test conditions ………………………………………....….. 33 Figure 20 General microstructure of YSZ composite processed at: a) 1300 °C, 30 MPa, 7 min, b) 1400 °C, 30 MPa, 7 min, c) 1500 °C, 50 MPa, 10 min………………..... 39 Figure 21 XRD plot for YSZ processed at: a) 1300 °C, 30 MPa, 7 min, b) 1400 °C, 30 MPa, 7 min, c) 1500 °C, 50 MPa, 10 min…………………………………….... 41 Figure 22 Representative SEM micrograph of 6.4 wt% CeO2 – YSZ after heat Treatment……………………………………………………………………….…. 43 Figure 23 SEM micrograph of 6.4 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min: a) general microstructure, b) matrix, c) Ce-rich particle………..... 44 Figure 24 Schematic of Cerium-rich particles in CeO2-YSZ composite……………………... 45 Figure 25 SEM micrograph of 12.7 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min: a) general microstructure, b) matrix, c) Ce-rich particles, d) microstructure after heat treatment……………………………………………... 47 Figure 26 SEM micrograph of 18.8 wt% CeO2 – YSZ composite sintered at 1300 °C, 30MPa, 5 min: a) general microstructure, b) matrix, c) Ce-rich particles, d) microstructure after heat treatment……………………………………………... 49 Figure 27 SEM micrograph of 18.8 wt% CeO2 – YSZ composite sintered at 1400 °C, 30 MPa, 7 min: a) general microstructure, b) matrix, c) Ce-rich particles………… 51 Figure 28 XRD plot for CeO2 –YSZ composites sintered at 1300°C, 30 MPa, 5 min: a) 6.4wt% CeO2, b) 12.7wt% CeO2, c) 18.8wt% CeO2, d) 18.8wt% CeO2 – YSZ composite sintered at 1400°C, 30 MPa, 7 min…………………………………….. 53 Figure 29 Line-scan analysis of Cerium-rich particle in 18.8 wt% CeO2-YSZ composite…... 54 xi Figure 30 Representative bimodal microstructure observed in La2O3 – YSZ composites: a) large and small grains indicated by arrows, b) small grain networks (dashed circles) containing voids (double arrows)…….…………….…. 57 Figure 31 Microstructure of 11.6 wt% La2O3 – YSZ composite sintered at 1400 °C, 30 MPa, 3 min: a) as-sintered matrix, b) large grains, c) small grains, d) small grains at 100,000x magnification………………………………………... 58 Figure 32 Microstructure of 11.6 wt% La2O3– YSZ composite sintered at 1500 °C, 50 MPa, 10 min: a) general microstructure, b) Lanthanum-rich particles, c) large grains, d) small grains………………………………………………….… 60 Figure 33 Microstructure of 21.2 wt% La2O3 – YSZ composite sintered at 1500 °C, 50 MPa, 10 min: a) general microstructure, b) Lanthanum-rich particles, c) large grains, d) small grains ………………………………………………….... 62 Figure 34 SEM-XEDS line-scan analysis of small grains in 21.6 wt% La2O3-YSZ…….…... 64 Figure 35 SEM-XEDS line-scan analysis of large particle in 21.6 wt% La2O3-YSZ……….. 65 Figure 36 Microstructure of 30.5 wt% La2O3 – YSZ composite sintered at 1500 °C, 50 MPa, 10 min: a) microstructure with primary (“A”) and secondary (“B”) matrices, b) large grains in primary matrix, c) large grains in secondary matrix, d) small grains in primary matrix………………………………………… 66 Figure 37 Large grains (bold arrow) observed in: a) primary matrix, b) secondary matrix of 30.5 wt% La2O3 – YSZ composite……………………………………... 68 Figure 38 Small grains (dashed arrows) observed in: a) primary matrix, b) secondary matrix of 30.5 wt% La2O3 – YSZ composite……………………………………... 68 Figure 39 SEM-XEDS line-scan analysis of secondary matrix in 30.5 wt% La2O3-YSZ composite showing lower Zr and higher La counts than in primary matrix……… 69 Figure 40 XRD plot for: a) 11.6 wt% La2O3 –YSZ sintered at 1400 °C, 30 MPa, 3 min, b) 11.6 wt% La2O3 –YSZ sintered at 150 0°C, 50 MPa, 10 min, c) 21.6 wt% La2O3 –YSZ sintered at 1500 °C, 50 MPa, 10 min, d) 30.5 wt% La2O3-YSZ intered at 1500 °C, 50 MPa, 10 min………………………………………………. 70 Figure 41 Representative bimodal microstructure observed in Nd2O3 – YSZ composites: a) large grains shown by bold arrow, b) small grains shown by dashed arrow….... 73 xii Figure 42 SEM micrograph of 12 wt% Nd2O3 – YSZ composite sintered at 1400 °C, 40 MPa, 5 min: a) as-sintered matrix, b) bimodal microstructure, c) large grains, d) small grains…………………………………………………………….. 74 Figure 43 SEM micrograph of 22.2 wt% Nd2O3 – YSZ composite sintered at 1400 °C, 40 MPa, 5 min: a) as-sintered matrix, b) bimodal microstructure, c) large grains, d) small grains…………………………………………………………….. 76 Figure 44 XRD plot for: a) 12 wt% Nd2O3 –YSZ sintered at 1400 °C, 40 MPa, 5 min, b) 22.2 wt% Nd2O3 –YSZ sintered at 1400 °C, 40 MPa, 5 min………………….. 77 Figure 45 Compressive creep response of YSZ sintered at 1400 °C, 30 MPa, 7 min……….. 79 Figure 46 SEM back-scattered micrographs of YSZ composite: a) before creep, b) after creep…………………………………………………………………………….… 81 Figure 47 SEM secondary-electron micrographs of YSZ: a) before creep, b) after creep..…. 82 Figure 48 Line-scans of particles observed in post-creep YSZ microstructure: a) analyzed site with labeled areas of interest, b) line scan of particle 1, c) line scan of particle 2………………………………………………………...…. 83 Figure 49 SEM micrographs of post-creep YSZ composite: a) crenellated cracked boundaries, b) trans-particle fracturing (arrows 1 to 3) and solid particle (arrow 4)…………………………………………………………………………... 84 Figure 50 Compressive creep response of 6.4 wt% CeO2-YSZ processed at 1300 °C, 30 MPa, 5 min…………………………………………………………………..… 86 Figure 51 SEM micrographs of general microstructure of 6.4 wt% CeO2-YSZ: a) before creep, b) after creep……………………………………………………………….. 87 Figure 52 SEM micrographs of Cerium-rich particles (arrows) in post creep 6.4 wt% CeO2-YSZ…………………………………………………………………………. 88 Figure 53 Cerium-rich particles in post creep 6.4 wt% CeO2-YSZ: a) SEM micrograph, b) screenshot of SEM-XEDS analysis software showing close up of textured vs. non textured particles………………………………………………………….. 89 Figure 54 Cerium-rich particles: a) trans-particle fracture, b) inter-particle fracture………... 90 Figure 55 Compressive creep response of 11.6 wt% La2O3-YSZ processed at 1400 °C, 30 MPa, 3 min…………………………………………………………………….. 92 Figure 56 SEM micrographs of 11.6 wt% La2O3-YSZ: a) before creep, b) after creep……... 93 xiii Figure 57 SEM micrographs of small grains in post-creep 11.6 wt% La2O3-YSZ: a) type I (without substructure), b) type II (with substructure)………………....… 94 Figure 58 SEM micrographs of post-creep 11.6 wt% La2O3-YSZ showing: a) grain boundary voids, b) voids (shown by arrows) located at boundaries and triple points, c) grain boundary interface with La-rich grains separated by voids …………………………..………………………………………………….. 96 Figure 59 Displacement of pistons vs. time plot obtained from register for compressive creep of YSZ at 1300°C and 627 N ……………………………………..…...….. 114 Figure 60 MS Excel spreadsheet with measured and calculated variables obtained during compressive creep test of YSZ at 1300°C and 627 N (Part 1 of 4)…….... 115 Figure 61 MS Excel sheet with calculated values for n and Q from a load and temperature jump………………………………………………………………... 116 Figure 62 MS Excel spreadsheet with measured and calculated variables obtained during compressive creep test of YSZ at 1300°C and 627 N (part 2 of 4)……..... 117 Figure 63 MS Excel spreadsheet with the measured and calculated variables obtained during compressive creep test of YSZ at 1300°C and 627 N (part 3 of 4)…….… 118 Figure 64 MS Excel spreadsheet with the measured and calculated variables obtained during compressive creep test of YSZ at 1300°C and 627 N (part 4 of 4)……..... 119 Figure 65 SEM-XEDS line-scan analysis of Ce-rich particle in 12.7wt% CeO2-YSZ composite……………………………………………………………………….... 120 Figure 66 SEM-XEDS line-scan analysis of Ce-rich particle in 18.8wt% CeO2-YSZ composite……………………………………………………………………….... 121 Figure 67 SEM micrograph of small grains in 12wt% Nd2O3 – YSZ at 100,000x magnification………………………………………………………………….…. 122 xiv List of Abbreviations AECL Atomic Energy of Canada Limited APS Air Plasma Spraying BSE Back-Scattered Electron CANDU Canada Deuterium Uranium CeO2 Cerium (IV) Oxide DC Direct Current GBS Grain Boundary Sliding HV Vicker’s Hardness KIC Fracture Toughness Kα X-Ray K-Alpha emission La2O3 Lanthanum Oxide LVDT Linear Variable Differential Transducer VLT Varying Load and Temperature Nd2O3 Neodymium Oxide NRC-CNBC National Research Council-Canadian Neutron Beam Center SCC Stress Corrosion Cracking SCWR Super Critical Water Reactor SE Secondary-Electron SEM-XEDS Scanning Electron Microscope-X-Ray Energy Dispersive Spectroscopy SOE Solid Oxide Electrolyte SOFC Solid Oxide Fuel Cell SPS Spark Plasma Sintering TBC Thermal Barrier Coating XRD X-Ray Diffraction YSZ 8mol% Yttria-Stabilized Zirconia ZAF Atomic Number (Z) Absorption (A) and Fluorescence (F) corrections xv List of Symbols ε Strain ̇ Strain Rate A Microstructure Constant D Diffusion Coefficient G Shear Modulus b Burger’s Vector K Boltzmann’s Constant T Temperature p Grain Size Exponent σ Stress n Stress Exponent Do Maximum Diffusion Coefficient Q Activation Energy R Universal Gas Constant Y Displacement of Pistons t Time Li Instantaneous Length of Sample Si Instantaneous Cross-section of Sample Lo Initial Length of Sample So Initial Cross-section of Sample xvi Acknowledgements First and foremost, I want to thank my supervisor Dr. Lukas Bichler for his relentless support and guidance for the past three years. Without your drive and knowledge, this research would not have been possible and I would not have been able to excel. Thank you for your patience, your advice, and your professional attitude that has helped me push through in moments of chaos. Audrey Siébert-Timmer: Thank you for these almost 6 years of friendship and for always bringing me back to reality when I thought there was no hope. I look forward to returning the support as you finish your own work, and definitely more sushi dinners. Karthik Akkiraju: Thank you for sharing your knowledge, your patience, and your reliable help throughout the entire research year and for sharing good laughs that others probably would never understand. I promise I will come visit you. Stephen Ipe: Thank you for your energy and enthusiasm, even though I know you would rather be watching a soccer match than polishing samples manually. I hope you find happiness in your future career and I hope we can see each other again one day. Mitchel van Hanegem: You thought I forgot about you, right? Thank you so much for all your help, I know at times it was a bit too much for you but you still followed through and never complained about a single thing. I had a lot of fun with the team socials and SEM Olympic marathons (even though sometimes you cheated!). I am sure you will have a fantastic future after graduation. Please stay in touch. Thank you, faculty members, who made a difference in my life and who continue to inspire me: Dr. Carolyn Labun, Laura Patterson, Dr. Ray Taheri, Dr. Vladan Prodanovic, Dr. Stephen O’Leary, Dr. Mina Hoorfar, Dr. Mondal, Dr. Murty and Dr. Spiro Yannacopoulos. To the laboratory gurus who saved me more than once: Russell LaMountain, Michelle Tofteland, David Arkinstall, Dean MacLellan and Durwin Bossy, thank you for your advice and excellent sense of humor. To the fantastic faculty at Universidad de Sevilla: Dr. Arturo Dominguez-Rodriguez, Dr. Miguel Castillo, Dr. Antonio Munoz, and to the lovely Jose, Ana Maria, Angelica, y Rosalia, muchas gracias por hacerme sentir parte de su familia. I also want to thank NSERC, AECL, and the NRC-CNBC for the expert advice and funding, which enabled me to collaborate with so many incredible individuals. Thank you Dr. Dimitry Sediako and Dr. Randy Fong for your help and always welcoming attitude. xvii Mariel Barrantes, Jessica Buriticá, Courtney Dean, Feng Haibo, Fabricio Bianchini, Arnaud Houriet, Hamid Reza Zareie, Juan David Torres Forero: you guys were my backbone and my support network, and a guaranteed source of awesome times. I am forever grateful for your friendship and your tough love. I wish you all a successful career in whichever path you follow. I will make sure to stay in touch. To my EWB family, thank you for being more than colleagues and for making me feel like there was always someone who I could approach regardless of the time of the year and hour of the day. You are all fantastic people and you have been there for me in the good, bad and the ugly moments of my studies. Thank you and good luck with your endeavors. Let’s stay crazy enough to think we can change this world. To the friends that have known me from the day I arrived to this country: Ian Roth, Milena Milenkovic, Maha Omar, Daniel González, Ana Clorio, Stephen Smithbower and Audrey Siébert Timmer. I would trust you with my life any day. Thank you again and again, you all know what you mean to me. I am happy to share the beginning of a new chapter in my life with all of you, and I am sure we will grow to see each other build happy lives and support each other along the way. You are simply the finest friends anyone could ask for. And to my multiple adoptive Canadian families, Dayna Li, Wendy and Howard Soon, Terri Flanagan, Denise Chan, and Cathie Roth. You have treated me like a daughter and helped me in my worst homesickness moments. Thank you. Gracias a mis padres, por enseñarme que no existe nada en este mundo que no pueda lograr. Por darme la major educación que una hija pudiera desear, y por demostrarme que no hay obstáculo a quién esta dispuesto a luchar. Gracias por su apoyo en todos los sentidos, sé que ah sido muy difícil y que todavía queda mucho camino por recorrer, pero pase lo que pase jamás dejare de quererlos y de admirarlos. Gracias por darme vida. Gracias a mis abuelo y a todos mis tíos; a mis primos, aunque sólo puedo verlos una vez al año, quiero que sepan que siempre pienso en ustedes y espero que puedan alcanzar todas sus metas. A mis hermanos, Kim y Luis, por no dejarme rendirme cuando no veía la salida. Espero poder haber sido la hermana que se merecían, aun cuándo no pude estar con ustedes durante todo el tiempo que quisiera. My family made me the person I am today. You allowed me to grow and stay strong despite the 3,750 km that separate us. Thank you. xviii Dedication To my brother and sister, Luis and Kimberly, and my mother Sandra. Luis and Kim: You are two of the strongest people I know. You have given me the strength I needed to get out of bed every morning. From the first time I arrived to Canada almost six years ago, I have strived to build a good future so I can share that with both of you. I hope I can make you as proud as you have made me. I love you more than any dedication could ever convey. Only you two could inspire me enough to work so hard for so long and still be able to smile. Este documento también está dedicado a mi madre. Mamá, gracias por tu apoyo y amor incondicional. Eres una de las mujeres más fuertes y valientes que conozco. Gracias por tu comprensión. Gracias por siempre estar ahí para mí. If I could have things my way, no family should ever need to be separated, and no child should ever have to struggle to attain a good education. I hope I can make this world a better place and that the next generation of Engineers has more readily access to the option of higher education. We live in a complex world, but there is nothing complex about wanting to do what is right. 1 Chapter 1: Introduction This chapter provides an overview of material development efforts to advance the Super Critical Water Nuclear Reactor technology. Candidate materials for use in pressure tubing components are presented and the methods for processing, characterization and creep testing are described. In the last decade, numerous international research groups [1] have started to investigate potential materials for use in the Generation IV (Gen. IV) nuclear reactors. These nuclear reactors are expected to meet the nuclear industry’s future demand of improved safety and sustainability in reactor design. One of the possible Gen. IV reactor designs is the “Super Critical Water Reactor” (SCWR). The SCWR was derived from the family of Canada Deuterium Uranium (CANDU) designs that were originally conceived by Atomic Energy of Canada Limited (AECL) in the 1950s. Canada’s current involvement includes research collaboration with government entities like AECL and the National Research Council-Canadian Neutron Beam Centre. Preliminary studies have shown that the SCWR offers benefits in terms of cost-savings, safety and increased sustainability. The SCWR would allow for production of energy at lower cost by as much as 35% compared to currently available reactors [2–4], as a result of high thermodynamical efficiencies (targeted between 45 % [1] and 50 % [5]). In addition to production of clean electricity, the nuclear waste generated by SCWR would decompose in a matter of decades instead of millennia as compared to current nuclear waste, which makes nuclear waste disposal in this reactor significantly more attractive. In order to accelerate the development of the SCWR technology, research groups must address the many challenges surrounding this reactor’s design, such as high temperatures, corrosive environment and neutron irradiation. Challenges to the design and material selection for different components in the SCWR derive from the hostile operating conditions involved with the use of water in a supercritical state. The operating conditions for the reactor’s components involve temperatures up to 600 C and pressure of 25 MPa. Materials would also be subject to radiation exposure up to 150dpa, which further contributes to an extremely aggressive environment. Therefore, 2 availability of materials showing outstanding creep resistance, corrosion resistance and irradiation stability are crucial for the future development of the SCWR technology. Various austenitic, ferritic, martensitic, and nickel based steels have been considered as potential materials for use in a SCWR [1], [2], [4–6]. In Buongiorno and MacDonald’s [2] progress report, for example, two austenitic alloys (304L, 316L) showed susceptibility to stress-corrosion cracking and pitting when tested in SCW above 400 C. Nickel-based alloys have also been found to experience pitting and oxidation, while oxidation of ferritic- martensitic steels has resulted in large weight gains [7]. Thus, the traditional steel alloys appear inadequate for the SCWR environment and other materials are being considered. One alternative to the steel alloy materials are Zirconium(Zr)-based superalloys, which can contain Titanium, Iron, Copper, and Molybdelum additives. These Zr-alloys have failed to overcome the issues of embrittlement and oxidation due to Zriconium’s high affinity for oxygen and absorption of hydrogen in the SCW environment [8–10]. As a result, research focus is shifting to ceramic materials. Specifically, Zirconia-based ceramic composites. Stabilized composites of Zirconia show excellent high-temeprature mechanical properties and superior creep resistance in comparison to steels. These composites are currently used for thermal barrier coatings (TBC) and solid oxygen electrolytes (SOE) [11– 14] where creep resistance is also important. Preliminary research has been carried out for potential uses of Zirconia and Zirconia-based ceramics within the SCWR; to list some examples, Zirconia ceramics could be implemented as an insulator in High Efficiency Channels (HEC), electric insulators or fuel coatings [15–17]. These components would require materials with low thermal conductivity, low neutron absorption and high creep resistance. Literature on creep resistance of Zirconia has shown that creep resistance can be achieved by suitable doping with rare earth oxides [12], [18–21]. The mechanisms responsible for enhanced creep resistance and the microstructure stability of the rare earth- doped cubic Zirconia are still subject of debate and data in published literature is limited. In the present research, novel composites based on the Zirconia matrix were developed using advanced fabrication technologies. Cubic Zirconia stabilized with 8 mol% Yttrium Oxide and additionally doped with rare earth oxides of Cerium Oxide, Lanthanum Oxide and Neodymium Oxide, were prepared by Spark Plasma Sintering (SPS) technique. 3 The composites were characterized with Scanning Electron Microscope –X-Ray Energy Dispersive Spectroscopy (SEM-XEDS), X-Ray Diffraction (XRD) and Optical Microscopy (OM). Also, compressive creep testing was carried out in the temperature range of 1200- 1330 C in an air atmosphere and values of the stress exponent and activation energy were determined. The possible creep mechanisms were then studied. This thesis document is structured as follows: Chapter 2 presents a literature review concerning the fabrication of rare earth oxide-doped Zirconia ceramics via the SPS technique, along with creep mechanisms observed in ceramics. Chapter 3 describes the experimental procedures used to fabricate the novel composites. Also, methods used to characterize materials and perform compressive creep tests are described. Chapter 4 discusses the results of SPS processing and compressive creep experiments, along with microstructure evaluation of the materials at various stages of testing. Chapter 5 provides the conclusions from this research. Chapter 6 presents topics and ideas suitable for future research. Chapter 7 provides the references used throughout this thesis. 4 Chapter 2: Literature Review This chapter presents a comprehensive literature review on the theory and general creep behaviour of ceramic materials. The effect of Cerium Oxide (CeO2), Lanthanum Oxide (La2O3) and Neodymium Oxide (Nd2O3) on the microstructure and creep resistance of Zirconia ceramics is discussed next. The chapter concludes with a discussion on the fabrication of Zirconia ceramics through Spark Plasma Sintering. 2.1 Introduction to Creep Resistance The term “creep” refers to a plastic deformation of a material over a long period of time. Creep can be measured in terms of strain ( ) or strain rate ( ̇ ), and is generally plotted against time or stress. A well-known approach to quantify creep resistance of a material is to derive strain rate from a strain vs. time curve. The strain rate experienced by a material can be related to several of the material’s properties, as indicated in Equation 1 below: ̇ ( ) ( ) Equation 1 Where: A = atomic structure constant D = diffusion coefficient G= shear modulus b = Burger’s vector k = Boltzmann constant (1.38·10 -23 J K -1 ) T = temperature p = grain size exponent σ = applied stress n = stress exponent 5 The diffusion coefficient, D, is defined as: ( ) Equation 2 Where: Do= maximum diffusion coefficient Q = activation energy R = universal gas constant Equations 1 and 2 can be used to estimate the strain rate, diffusion coefficient, stress exponent, grain size exponent and activation energy from experimental measurements obtained during a creep test. In turn, these parameters assist in identifying the microstructural mechanisms involved in the creep response of a material. Combining Equations 1 and 2 allows for a simplified approach of quantifying the strain rate and estimation of the parameters stress exponent and activation energy. The simplified expression is shown in Equation 3 ( a detailed derivation of Equation 3 is presented in reference 17) : ̇ ( ) Equation 3 The stress exponent, n, helps to identify the type of creep mechanism at a particular temperature and pressure, while the activation energy, Q, is the amount of energy needed to activate this mechanism. The stress exponent can also be calculated from the slope in a logarithmic plot of strain rate vs. stress, while Q can also be calculated from the slope of a logarithmic plot of strain rate vs. inverse temperature. Extensive experimental data, as seen in Figure 1, has enabled estimation of n and Q for materials in relation to their properties (e.g. grain size) [22–24]. 6 Figure 1 Logartithmic plots used to estimate: a) n, b) Q, c) p. [25], [26] The determination of the active creep mechanisms during deformation cannot be done based on one variable. Ideally, the stress exponent and activation energy, and in most cases the grain size exponent, p, should be considered and evaluated in conjunction with micrographs, grain size measurements, and chemical phase analysis. For example for creep of Zirconia ceramics, it is usually accepted that a stress exponent of n~2 is related to grain boundary sliding, while n~1 is related to lattice diffusion mechanisms [19], [22], [26–28]. In the present work, however, only the stress exponent and activation energy parameters were investigated. Despite the efforts to correlate experimental data to deformation mechanisms, conclusive identification of creep mechanisms in ceramic materials is hindered by the dynamic nature of creep. The experimentally measured values for n and Q tend to span a wide range of values with temperature, pressure, and time, suggesting superposition or simultaneous regimes being active [18], [25], [29], [30]. From Cannon and Langdon’s [19], [20] and Jimenez-Melendo’s [24] extensive reviews on the creep of ceramics, a general trend 7 was defined were an exponent of n~1-3 was related to grain boundary sliding, while an exponent of n~3-5 corresponds to lattice diffusion mechanisms. The most frequent mechanisms arising from compressive creep experiments are described next. 2.2 Creep Mechanisms Creep mechanisms help identify the nature of plastic deformation of materials, in relation to material properties and ambient conditions. In general, the creep mechanisms are classified as follows [19]: 1. Diffusion (Lifshitz) Creep i) Lattice Diffusion (Nabarro-Herring Creep) ii) Grain Boundary Diffusion (Cobble Creep) 2. Grain Boundary Sliding (Rachinger) Creep 3. Dislocation Slip/Climb Creep Diffusion or Lifshitz creep refers to a mass or vacancy transport either through the lattice (Nabarro-Herring) or along the grain boundaries (Cobble), as shown in Figure 2, and results in elongation of the grains due to tensile stress. The mass transport can be of anions, cations, or dislocations. Nabarro-Herring creep occurs more often than Cobble creep, because there are more paths for the atoms to diffuse; however, it is a slower process than Cobble creep. Both types of diffusion creep are accelerated with increasing ambient temperature. 8 Figure 2 Lattice diffusion mechanisms. [31] Rachinger creep, or grain boundary sliding (GBS), occurs as a result of shear stresses acting along the grain boundaries and results in the formation of voids along the grain boundaries. As shown in Figure 3, the grain boundaries move with respect to the applied stress (indicated by double arrows), which may result in the nucleation of a crack. Figure 3 Grain boundary sliding schematic showing void formation. [31] 9 The specific composition of the ceramic must be taken into account, as research has shown that precipitate diffusion, glassy phase formation, and strain hardening due to certain ions also affect the creep rate [27], [28], [31], [32]. For example, the presence of ion segregation at the grain boundaries was seen to slow down dislocation movement [34]. In contrast, the presence of a glassy phase on the grain boundaries was related to grain boundary-diffusion [12], [18], and was seen to contribute to superplasticity [34]. As dislocations in a material move, they encounter obstacles at the grain boundaries and therefore i) slow down [30] or ii) are pinned at that boundary. Both events result in slowing down of dislocations [35] and overall decrease of the strain experienced by the sample [22], [36]. Other research has shown that certain compositions of rare-earth doped Zirconia ceramics can cause a sample to undergo strain hardening [30] during loading, which further contributes to creep resistance in the material. 2.1 Zirconia Ceramics Zirconia is a polymorphous material, which can exhibit monoclinic, tetragonal or cubic crystal structure, depending on the temperature. The three phases of Zirconia are depicted in Figure 4. Figure 4 Crystal structure of Zirconia. [37] At room temperature Zirconia has a monoclinic structure. As temperature is increased (~1170 C) [38] a phase transformation is induced and the tetragonal phase forms. With further temperature increase (~2716 C) [38] cubic structure forms. The lattice angle in the monoclinic lattice is 98 °, while in the tetragonal and cubic lattices the angle is close to 90 ° 10 [21]. Further, the lattice parameters a and b are equal only in the cubic lattice. Therefore, as Zirconia undergoes phase transformations, lattice geometry changes lead to significant volumetric lattice changes, causing material cracking and residual stresses [21], [39], [40]. The tetragonal and cubic phases of Zirconia can be stabilized to enable them to exist at room temperature. Stabilization is achieved with the addition of divalent or trivalent oxides such as MgO, Y2O3 and CeO2. With respect to creep behaviour of Zirconia-based ceramics, the majority of literature is focused on fine-grained, tetragonal polycrystalline Zirconia, due to its promising superplastic characteristics. Jiménez-Melendo et al. [24] carried out an extensive literature review on creep behaviour of Zirconia-based ceramics, and their investigation showed a lack of information on cubic Zirconia given that most of the studies reviewed were concentrated on superplastic behaviour. Creep behaviour of polycrystalline cubic Zirconia has only been explored by a few [12], [18], [19], [28], [41]. Doping of cubic Zirconia by rare earth oxides has shown a great promise in improving the mechanical properties at elevated temperatures [12], [13], [18], [20]. However, there is a paucity of literature on this subject. The following sections provide an overview of available literature on doping of Zirconia by Ceria, Lanthana and Neodymia. 2.2 Addition of Cerium Oxide (CeO2) to Zirconia Ceria-stabilized Zirconia ceramics have been used in TBCs and in solid oxide fuel cells (SOFC) [42], [43] due to their excellent thermo-mechanical properties [33]. The solubility of Cerium in Zirconia is restricted due to the large difference in ionic radii. Available phase diagrams of Ceria-Zirconia and Ceria-Yttria-Zirconia systems are provided in Figure 5. In Figure 5a, the phases of cubic solid solution (css), tetragonal solid solution (tss) and monoclinic solid solution (mss) have been indicated. 11 Figure 5 Phase diagrams: a) Ceria-Zirconia binary system, b) Ceria-Yttria-Zirconia system at 1400 °C in Zirconia-rich region.[44] The location of solubility lines in the CeO2 –ZrO2 phase diagram (Figure 5a) remains subject of debate, as contradicting information suggests values for the solubility of CeO2 in ZrO2 as low as <5 mol% [45] or as high as 40 mol% [46], [47] in the temperature range of 1300 to 1400 C. Boutz et al. [48] studied the fracture toughness, ageing resistance and sinter forging of 2, 4, 6 and 8 mol% ceria-doped 4 mol% Yttria-stabilized Zirconia (YSZ). The composites were prepared through pressure-less and pressure-assisted conventional sintering at temperatures of 1150-1550°C, pressure of 80-100 MPa for 10 hrs. The grain size of as- sintered composites was ~0.21 m, with relative densities ranging between 92-99 %. Boutz et al. found that composites sintered with pressure assistance had higher fracture toughness than those obtained via pressure-less sintering (8-9 MPa m 1/2 versus 5-7 MPa m 1/2 ). The study also found that the stability of the tetragonal phase improved with decreasing of grain size and increase in ceria concentration. There was no clear trend observed between fracture toughness and grain size, while fracture toughness values were generally lower for composites with increased ceria content. The fracture toughness of the composites was also seen to increase as porosity decreased. These findings suggest that sintering pressure plays an 12 important role on the consolidation and the ensuing mechanical properties of CeO2-YSZ composites. Tsipas [43] studied the phase stability of a commercial Yttria-stabilized Zirconia coating with ~25 wt% cerium oxide and ~2.4 wt% yttrium oxide. The ceria-containing coating was applied to a metallic surface through air plasma spraying (APS). Subsequently, the coating was annealed at 500 °C (holding time not specified) in vacuum. Tsipas [43] reported that the Ce 4+ cations were able to replace the Zr 4+ cations in the tetragonal Zirconia structure. However, since the Ce 4+ is a larger ion than the Zr 4+ , the lattice experienced strain, which decreased the tetragonality of the crystal lattice. Further, Ce 4+ reduction to Ce 3+ was associated with a volume change and decomposition of the tetragonal phase. The presence of Ce 3+ and Y 3+ in Zirconia composites was also discussed by Matsuzawa and Horibe [49] and de Bernardi-Martín et al. [33]. The presence of Ce 3+ was seen to introduce oxygen vacancies in the crystal lattice in order to retain electrical balance; these vacancies enabled the mobility of ions within the lattice in the presence of applied stress, resulting in higher creep rates. In the case of Y 3+ , ion segregation to the grain boundaries occurred as the material underwent deformation, and resulted in the formation of an “electric charged layer” at the grain boundaries, which was claimed to have a creep-suppressing effect. de Bernardi-Martín et al. [32], [33] also studied compressive creep of 12 mol% cerium- tetragonal Zirconia. Cold isostatic pressing of the raw powders followed by sintering at 1450- 1500°C was carried out. In the first study, de Bernardi-Martín et al. [32] performed compressive creep between 1200-1250 °C with pressure ranging between 84-122 MPa. The measured stress exponent was 1 ± 0.3 and the activation energy was 300 ± 50 kJ/mol. Although the value of n was representative of diffusional creep, the authors suggested that GBS took place since no grain growth was observed after deformation. This led the authors to suggest GBS was the controlling creep mechanism, but further investigation was deemed necessary to verify this result. In the second study [33], de Bernardi-Martín et al. carried out constant pressure and temperature creep tests. The first creep experiment was carried out between 1200-1300 C at a constant pressure of 125 MPa; the second creep experiment was carried out at a constant temperature of 1300 C and pressures ranging between 28-80 MPa. The measured stress 13 exponent was 1 ± 0.2 and activation energy 550 ± 50 kJ/mol. Voids at the triple points of the grain boundaries were reported, and it was suggested that creep of Ce-ZrO2 materials may be controlled by grain boundary diffusion. To enhance the creep resistance, the authors proposed dispersion of obstacles along the boundaries to slow down the relative sliding and subsequent cavitation at the grain triple points. 2.3 Addition of Lanthanum Oxide (La2O3) to Zirconia Lanthana- Zirconia ceramic composites are being developed for TBCs as well as SOE due to their low thermal conductivity and high temperature stability [43], [50–54]. The phase diagram for the Lanthana-Zirconia system is shown in Figure 6. Figure 6 Phase diagram for Lanthana-Zirconia system. [55] According to Bastside et al. [56], the maximum solubility of tetragonal phase in this system is 0.75 mol% La2O3 between 1100 and 1500 C; at room temperature, the authors expect the system is biphasic (monoclinic and pyrochloric solid solutions) between 0.5 and 20 mol% La2O3. However, other studies [50] suggest that more consistent experimental data is necessary to confirm the solubility limits of this system. 14 Bastide et al. [56] prepared composites of 0.5-12 mol% Lanthana in Zirconia matrix via conventional sintering. Raw powders were pressed isostatically at 480 MPa, followed by sintering between 1200-1600 C at a rate of 100 C/min for 1.5-2 hr. They observed that the addition of Lanthana had a strong grain refinement effect when composites were sintered above 1450 °C. For instance the grain size reduced from 8 μm (0 mol% La2O3) to as low as 0.2 μm with 10 mol% La2O3 addition. Further, XRD work revealed that the tetragonal phase appeared in the compacts around 3 mol%, while the pyrochlore structure appeared at 8 mol%, and a combination of tetragonal/pyrochlore structures was present at La2O3 concentrations in excess of 10 mol%. Singh et al. [57] investigated the stabilization of the cubic and monoclinic Zirconia phase by addition of 1-24 mol% Lanthana. The composition was prepared using liquid-mixed technique and compacts were calcined at 800 °C for 4 hrs. XRD analysis revealed that the cubic phase was successfully stabilized with the addition of 4-24 mol% Lanthana. Wang et al. [50] investigated the Lanthana-Zirconia system with 25 and 75 mol% La2O3 added to Zirconia. The composites were sintered between 1400 and 1700 °C, with heating and cooling rates of 10 °C/min. XRD analysis and SEM backscattering confirmed the presence of pyrochlore and monoclinic structures in the 25 mol% Lanthana-Zirconia composites. The 75 mol% Lanthana-Zirconia composite consisted entirely of pyrochlore and the La(OH)3 structure. Tsipas [43] investigated the phase stability of YSZ coating containing 6.58 wt% (~2.7 mol%) Lanthana. The coating was heat treated at 1200-1400°C and was cooled at a rate of 30 °C/min. The coating was originally composed of tetragonal YSZ only; after heat treating, however, the tetragonal phase decomposed into a combination of tetragonal, cubic and pyrochloric phases of YSZ. Several researchers have investigated the 33 mol% Lanthana-Zirconia system, also known as Lanthanum Zirconate (La2Zr2O7) [52–54], [58], [59]. Wang et al. [54] investigated the effect of sintering temperature on the grain size of Lanthanum Zirconate composite. Samples were compacted at 220 MPa in air, followed by sintering at 1500 °C at a rate of 5 °C/min for 6 hrs. An increase of grain size was observed with increasing sintering temperature, with the smallest grain size of ~15 nm obtained at 500 °C and the largest grain size of ~65 nm obtained at 1100 °C. Similar trends were reported by other researchers, and 15 suggest that the grain size in Lanthana-Zirconia composites has a strong dependency on the sintering temperature even with very high Lanthanum contents. 2.4 Addition of Neodymium Oxide (Nd2O3) to Zirconia The available literature on polycrystalline Neodymia-Zirconia composites is very limited. The Neodymia-Zirconia composites are expected to offer low thermal conductivity, high melting points, and a complex crystallographic structure, which would make this composite useful for TBCs, solid electrolytes, and diode pumped optical solid state lasers [51], [60], [58]. A phase diagram for the Neodymia-Zirconia system is shown in Figure 7. Figure 7 Phase diagram for Neodymia-Zirconia system developed by Wang et al. [50]. Surkov et al. [62] investigated the sinterability of 0.5-10 mol% neodymium-stabilized Zirconia composites through dilatometric studies and XRD analysis. The composites were sintered at 1750 C and a pressure of ~49 MPa at varying heating rates. High porosity and high water absorption were seen with increasing neodymium content and lower sintering temperatures. XRD analysis revealed that monoclinic Zirconia phase was present for 16 neodymium concentrations of as low as 0.5-2 mol%. For the 6 mol% Neodymia-Zirconia composite, the analysis confirmed the tetragonal phase of Zirconia along with some monoclinic phase were present. At neodymium concentrations above 10 mol%, neodymium addition stabilized the cubic and tetragonal Zirconia solid solution. Xu et al. [14] investigated the relationship between microstructure, fracture toughness, Vicker’s hardness and sintering of Neodymia-Zirconia composites. A colloidal processing technique was used to prepare 1.5–2 mol% Nd2O3–1.5 mol% Y2O3-coated monoclinic ZrO2 powder; this powder was sintered in air at 1450 C for 1-5 hrs. Xu et al. obtained fully dense compacts after sintering for up to 3 hrs. Sintering for 3 hrs induced a monoclinic phase transformation, which was accompanied by a volume change (3-4%) and consequent material cracking. For the 1.5 mol% Nd2O3–1.5 mol%Y2O3 composite, the value of KIC was 11.14 MPa m 1/2 , representing a 5-8 % increase in comparison to YSZ [21]. Sintering for longer than 3 hrs resulted in undesirable grain growth, and a spontaneous tetragonal-to-monoclinic transformation, which created microcracks and contributed to the loss of strength. Salehi et al. [63] investigated the densification, Vicker’s hardness, fracture toughness and low temperature degradation of monoclinic ZrO2 nanopowder coated with a blend of 1–2 mol% Nd2O3 + 1 mol% Y2O3. The powders were densified through Spark Plasma Sintering at 1450 C and 60 MPa with a holding time of 3 min. Fully dense compacts were obtained for both concentrations of neodymium. Increasing the neodymium content from 1 to 2 mol% reduced the grain size from ~0.4 μm to ~0.25 μm, reduced KIC from 7.4 to 5.9 MPa m 1/2 , and increased HV10 from 1076 to 1117 GPa. The effect on KIC was attributed to the presence of over-stabilized tetragonal phase, thus hindering the transformation toughening mechanism and lowering the fracture toughness. The preparation of composites discussed thus far involved fabrication using one of the most advanced sintering process—the Spark Plasma Sintering (SPS) technique. The following section describes the fundamental characteristics of SPS processing. 2.5 Overview of Spark Plasma Sintering of YSZ Spark plasma sintering (SPS) process was developed in the 1960’s [64]. Due to patent restrictions, commercialization of SPS was delayed [65]. Nowadays, SPS has been recognized for its ability to prepare material compositions that cannot be achieved through 17 conventional sintering. SPS can be used for consolidation of metals, ceramics, polymers, functionally graded materials and nanocomposites [64], [66]. The SPS process is capable of sintering fully dense samples, while retaining the nano-grain structure of the original powders. The SPS process begins with the loading of raw powders into a circular sintering graphite die (optional use of binding agents), and graphite sheets are placed on the top and bottom ends of the die to create a seal in the die. Graphite is the typical material used for the die and sheets due to its high temperature strength, density, and relatively low cost [65]. The necessary heating and cooling rates, holding temperature, holding time, applied pressure and atmosphere are programmed into the SPS machine [65]. Typical heating rates range between 100 and 200 C /min [67–69], while the holding temperature and time as well as atmosphere are strongly dependent on the material being sintered. The cooling rate may be programmed into some SPS machines, although it is also common for samples to be cooled naturally. The combination of all the parameters used for sintering a material are usually called a “cycle”. After the sample is cooled down, it is stripped out of the die manually and is ready for use. A typical SPS machine (Figure 8) consists of a sintering chamber, sintering die, top and bottom punches, thermocouples, and the DC pulse generator. A pulse-formed electric discharge is generated and provides Joule heating to facilitate fusion of powder particles inside the sintering die [64], [68]. 18 Figure 8 SPS apparatus schematic. [70] Literature on preparation of Zirconia composites using SPS is very limited. Bernard- Granger and Guizard [67] performed systematic SPS experiments of 3 mol% Yttria- stabilized tetragonal Zirconia at varying temperatures (950-1200 °C) and holding times (5- 180 min). The following general trends were observed for SPS processed Zirconia composites:  Relative density increased with increasing sintering temperature, regardless of sintering time  The grain size increased as sintering time increased  A relative density as high as 98 % can be obtained. Bangchao et al. [68] sintered 8 mol% Yttria-stabilized Zirconia, starting powder with particle size 0.1 m, using conventional sintering (CS) and the SPS technique. CS was carried out using uniaxial compaction at 140 MPa and subsequent sintering at 1450°C for 4 hrs. SPS sintering was completed using a 15 mm diameter graphite die inside a sintering chamber evacuated to 5 Pa; the ram pressure of 28 MPa and sinter temperature of 1350°C 19 were held for 10 min. A heating rate of 160 °C/min was followed by natural cooling. The compacts were annealed in air at 1000 °C for 2 hrs to remove carbon contamination. The obtained microstructures from SPS and CS compacts are shown in Figure 9. Figure 9 SEM micrographs of cubic Zirconia prepared with: a) SPS, b) conventional sintering. [68]. Clearly, SPS processing enabled fabrication of fine-grained material. The highest relative density for conventional sintered compacts was 97 %, while for the SPS sintered compacts it was 99 %. The SPS sintered sample also had a finer grain size (0.154 m) than the conventionally sintered sample (>1 m). In 2009, Biswas [66] processed 8 mol% Scandia-stabilized Zirconia by SPS, shown in Figure 10. Figure 10 SEM micrograph of Scandia-stabilized Zirconia prepared with SPS. [66]. 20 The precursor powder (particle size ~17 nm) was sintered at 1350 °C for 5 min with a heating rate of 500 °C/minute in a vacuum chamber, with a compaction pressure of 60 MPa. The author reported successful preparation of high-density compacts with an average grain size of <100 nm. 2.6 Chapter Summary Conducting a literature review on the fabrication of YSZ composites revealed that the composite grain size increases with increasing sintering temperature, regardless of the oxide addition. In the case of CeO2 and La2O3 addition, the grain size decreases with increasing dopant concentration. Phase stabilization appeared to be a challenge for the majority of the studied composites. Limited literature is available on SPS processing of polycrystalline Zirconia ceramics. The studies reviewed in this section suggest that densification improves with an increase of sintering temperature, and grain size increases with sintering time. Also, rare-earth addition was seen to improve densification of Zirconia ceramics. 21 Chapter 3: Experimental Procedure This chapter describes the materials, SPS fabrication process and evaluation techniques for the composites studied in this research. 3.1 Precursor Powders Precursor powders of 8 mol% Yttria-stabilized Zirconia as well as rare earth oxides (CeO2, La2O3 and Nd2O3) were obtained from Inframat Advanced Materials (CT, USA). MSDS sheets from Inframat showed that the powders were produced via the wet chemical co-precipitation method and had a purity of 99.9 % + (metal basis excluding Hf). The powder composition and particle size of each powder (obtained by supplier through BET) is provided in Table 1. SEM micrographs of the powders are presented in Figure 11. The SEM images reveal that the raw powders had a tendency to form agglomerates (or clusters). This clustering could potentially make preparation of homogeneous as-sintered materials difficult. As a result, special powder blending procedure was implemented (as discussed in Section 3.2) to ensure uniform chemistry of the final powder blends. SEM- XEDS revealed traces of Cl and Hf in the precursor powders, which likely originated from the powder manufacturing process. Table 1 Chemical composition of precursor powders. Precursor Powder Concentration (wt%) Particle Size (nm) Zr O Y Ce La Nd Cl Hf YSZ 60.4 28.7 8.8 - - - 0.7 1.3 500 CeO2 - 17.8 - 81.2 - - 0.8 0.1 50-80 La2O3 - 24.1 - - 75.9 - - - ~80 Nd2O3 - 17.6 - - - 76.7 5.5 <0.1 ~80 22 Figure 11 SEM micrographs of precursor powders: a) YSZ, b) CeO2, c) La2O3, d) Nd2O3. 3.2 Precursor Powder Blending Blending of the precursor powder of CeO2 with YSZ matrix was carried out at the Indian Institute of Technology Madras (Chennai, India), by the research group under Dr. B. S. Murty. The powders were first weighed using an electronic balance and then placed in a Ball-milling vial. Mixing was carried out in a planetary ball mill for 15 min at 300 rpm without milling balls or liquid binders. Mixing of precursor powders of La2O3 and Nd2O3 with YSZ matrix was carried out at the University of British Columbia. Precursor powders were placed in a beaker with alcohol and stirred automatically for 4 min using a rotating 23 magnet on a hot plate. The mixture was allowed to settle for 24 hrs at room temperature and, once dry, powders were manually ground using a porcelain pestle for 20 min. Composite powders of various doping levels were prepared as outlined in Table 2. Table 2 Levels of doping of YSZ with rare earth oxides. Doping Level CeO2 La2O3 Nd2O3 wt% mol% wt% mol% wt% mol% 1 6.40 4.96 11.60 5.02 12.00 5.05 2 12.7 9.99 21.6 10.0 22.2 10.0 3 18.8 15.0 30.5 15.0 - - 3.3 Spark Plasma Sintering Due to the limitation in the availability of SPS machines in Canada, SPS processing was carried out at two locations. YSZ and CeO2 –YSZ composites were prepared at the Indian Institute of Technology Madras (Chennai, India), while La2O3 –YSZ and Nd2O3 –YSZ composites were prepared at Quad City Manufacturing Lab (Rock Island, USA), by qualified technicians in each institution. 3.3.1 Sintering of YSZ and CeO2-YSZ Composites The YSZ and CeO2 –YSZ composites were sintered using Dr. Sinter (SPS-625) Spark Plasma Sintering machine. The powders were placed in a 20 mm diameter and 40 mm tall graphite die. The exact processing parameters for each composite are listed in Table 3. Table 3 SPS processing parameters for YSZ and CeO2-YSZ composites. Composite Sintering Temperature (°C) Sintering Pressure (MPa) Sintering Time (min) YSZ 1300 30 7 1400 30 7 1500 50 10 6.4wt% CeO2 - YSZ 1300 30 5 12.7wt% CeO2 - YSZ 1300 30 5 18.8wt% CeO2 - YSZ 1300 30 5 1400 30 7 For all samples, the heating rate was 100 ºC/min. After the desired sintering time elapsed, the vacuum was released (leak valve released and pumps disengaged). This release was done at least 5 min after the sintering die reached a temperature of 600 °C. The 24 composites were allowed to cool in the furnace naturally since the utilized machine did not allow direct control of the cooling rate. Representative samples of as-sintered discs of the YSZ and CeO2-YSZ composites are shown in Figure 12 and Figure 13. Figure 12 Representative samples of as-sintered: a) YSZ sintered at 1300 °C, 30 MPa, 7 min, b) YSZ sintered at 1400 °C, 30 MPa, 7 min, c) sintered at 1500 °C, 50 MPa, 10 min. 25 Figure 13 Representative samples of as-sintered: a) 6.4 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min, b) 12.7 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min, c) 18.8 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min, d) 18.8 wt% CeO2 – YSZ composite sintered at 1400 °C, 30 MPa, 7 min. 3.3.2 Sintering of La2O3-YSZ and Nd2O3-YSZ Composites Precursor powders of blends of La2O3 –YSZ and Nd2O3 –YSZ were sintered using a Thermal Technology Spark Plasma Sintering machine (10-3 model) at QCML, USA. The powders were placed in a 19 mm diameter graphite die with a grafoil lining. A load of 5 MPa was applied prior to sintering of the powders and the sintering chamber was pumped down to 26 5x10 -2 Torr. The sintering chamber was then heated at 200 °C/min until the temperature reached 600 °C; this temperature was sustained for 10 s followed by further heating at the same rate until the desired sintering temperature was reached. After the desired pressure was applied for the duration of the specified sintering time, the composites were allowed to cool in the furnace naturally. The SPS processing parameters for each composite are listed in Table 4 and representative as-sintered composites are shown in Figure 14 and Figure 15. Table 4 SPS processing parameters for La2O3-YSZ and Nd2O3-YSZ composites. Composite Temperature (°C) Pressure (MPa) Time (min) 11.6 wt% La2O3 - YSZ 1400 30 3 1500 50 10 21.6 wt% La2O3 - YSZ 1500 50 10 30.5 wt% La2O3 - YSZ 1500 50 10 12 wt% Nd2O3 - YSZ 1400 40 5 22.2 wt% Nd2O3 - YSZ 1400 40 5 27 Figure 14 Representative samples of as-sintered: a) 11.6 wt% La2O3 –YSZ composite sintered at 1400 °C, 30 MPa, 3 min, b) 11.6 wt% La2O3 –YSZ composite sintered at 1500 °C 50 MPa, 10 min, c) 21.6 wt% La2O3 –YSZ composite sintered at 1500 °C 50 MPa, 10 min, d) 30.5 wt% La2O3-YSZ composite sintered at 1500 °C 50 MPa, 10 min. 28 Figure 15 Representative samples of as-sintered: a) 12 wt% Nd2O3 –YSZ composite sintered at 1400 °C, 40 MPa, 5 min, b) 22.2 wt% Nd2O3 –YSZ sintered at 1400 °C, 40 MPa, 5 min. 3.4 Polishing, Thermal Etching and Density Measurement In order to prepare samples for metallographic analysis, surface polishing of the composites was necessary. A sample preparation procedure outlined in Table 5 was followed for all composites. To assist with grain size analysis, polished samples were thermally etched in air at 50 ºC below their respective sintering temperature. The thermal treatment, programmed into the furnace (UAF-Lenton Chamber Furnace, MHI High Temperature Furnace) control unit, was carried out for 1 hr with a heating rate of 10 ºC/min and a cooling rate of 20 ºC/min. The specific gravity of the composites was measured using the Archimedes’ method. An electronic balance with a beaker with water (at room temperature) was used. The relative density was calculated by dividing the measured density by the theoretical density and was expressed as a percentile. 29 Table 5 Polishing procedure for microstructure evaluation. Step Grit/Paste Time Coolant 1 SiC 240 Until flat Water 2 SiC 600 10 min 3 SiC 1200 20 min 4 Diamond 9 μm 30-60 min metaDi® 5 Diamond 6 μm 30-60 min 6 Diamond 3 μm ~30 min 3.5 Microscopy and XRD Optical microscopy and scanning electron microscopy was carried on composites at various stages of testing. A Zeiss Axiovert optical microscope with bright field and DIC modes was used for macro imaging. Tescan MIRA3 XMU scanning electron microscope (beam size of 7 nm) with Oxford X-max XEDS detector (without ZAF correction) were used for high resolution electron imaging and chemical analysis. Buehler Omnimet image analysis software was used to measure the area fraction of secondary phases, grain sizes and grain morphology. Also, the software was used to measure the grain size via the linear intercept method. X’Pert Pro (PANalytical) X-Ray Diffractometer with Bragg-Brentano set-up, equipped with X’Celeratirm was used for XRD analysis. Diffraction peak measurements were performed in a range of 20-90º using 0.02º step size with a 20s time per step along with Cu- K radiation (45 kV, 30 mA). 3.6 Compressive Creep Experiments Compressive creep testing was carried out at the Department of Physics of Condensed Matter at the University of Sevilla, Sevilla, Spain, by the author. The following text provides an overview of the instruments used, measurements performed and subsequent calculations required to determine the stress exponent and activation energy of composite materials investigated in this research. 3.6.1 Sample Preparation Test specimens were sectioned into parallelipeds with dimensions of ~2.5 x 2.5 x 5 mm 3 using a diamond blade saw and a Struers® rectifier. An example of a prepared paralleliped 30 ready for testing is shown in Figure 16 (representative samples shown in the figure are Alumina composite). Figure 16 Samples for creep testing: a) Sample location within the as-sintered disc, b) dimensions of creep specimen. 3.6.2 Creep Apparatus Compressive creep testing was performed using the patent-pending apparatus shown in Figure 17. The apparatus consists of a compression chamber, weight holder, programming unit, analogue length gauge and an analog register. The interior of the compressive chamber, illustrated in Figure 17b, included a pair of Silicon Carbide (SiC) pistons, tungsten resistive sheets, and a moving top crosshead. Prior to performing a test, the compression chamber was completely sealed and was purged with Argon. Then, electrical resistance heaters were turned on. The temperature inside the chamber was monitored using thermocouples. The programming unit controlled the temperature, heating rate, vacuum level, and the dwell time (i.e. the time that a specific temperature was maintained during a test). Sample deformation and piston displacement were measured using a linear variable differential transducer (LVDT) (TESATRONIC TT60®). The piston displacement was recorded as a function of time on a paper roll register. The required load for testing was imposed on the sample by placing weights onto a platform of the machine. Thus, the load was recorded in Newtons. When calculating the stress imposed on the sample, the instantaneous cross-section of the sample was used. Hence, the stress values represent the true stress. 31 Figure 17 Compressive creep apparatus: a) front view of the machine, b) schematic of the compression chamber. 3.6.3 Creep Testing Procedure Two types of compressive creep tests were performed for each sample: i) a uniaxial compression test at a constant crosshead speed, and ii) a varying load and varying temperature test. The first test was used to estimate the minimal load needed to achieve plastic deformation. The second test was performed to enable calculation of n and Q. 3.6.3.1 Constant Crosshead Speed Test The objective of the constant crosshead speed test was to obtain a force vs. strain plot in order to estimate the onset of plastic deformation in the composite. Figure 18 shows a typical stress versus strain graph obtained from the experiment. 32 Figure 18 Stress vs. strain curve for YSZ at 1300ºC and a deformation rate of 1μm/min. The test procedure was developed by the research group of Prof. Dominguez-Rodriguez, and was performed by Dr. Castillo, a PDF in the group. First, the samples were glued to two SiC pellets and placed between the pistons in the compression chamber (Figure 17b). Once the sample was positioned, the chamber’s rubber lining was coated with grease (Dow Corning® high vacuum grease) to ensure a proper seal. Then, the vacuum was established (~10 -3 mbar). Thereafter, the controller was programmed to heat up the chamber to a temperature of ~500 ºC at ~100 ºC/min. Once at this temperature, Argon washes were performed and high vacuum (~10 -7 mbar) was established. The desired test temperature (e.g. 1300 ºC) was programmed into the controller and the system was allowed to reach a stable temperature condition before the necessary weights were added to the apparatus in order to initiate deformation. The stress vs strain diagram was then generated and examined for the onset of plastic deformation. For example, the plot in Figure 18 shows that a stress of ~90 MPa was required for plastic deformation of a particular composite. For this sample the subsequent test carried under varying load and temperature conditions would begin at an initial stress of ~90 MPa. 33 3.6.3.2 Varying Load and Temperature Test The objective of the varying load and temperature (VLT) test was to measure the strain and strain-rate during deformation in order to calculate n and Q parameters. During the test, varying forces (or loads) were exerted on the sample held at a constant temperature, followed by constant force test with a varying temperature. A schematic of the VLT test profile is illustrated in Figure 19. Figure 19 Schematic of the VLT test conditions. As the sample compressed, the displacement of the pistons was measured using a LVDT. The test continued until the onset of the secondary creep (i.e. constant slope on the displacement vs. time plot) was reached. Thereafter, the load or temperature was varied and the test repeated. Usually, five load jumps were performed during a single test. Similarly to varying load, the temperature was varied in a temperature jump. The maximum temperature reached during any of the creep tests was at least 50 ºC below the sample’s sintering temperature in order to avoid any microstructural instability (such as grain growth), which could compromise the experimental results. A summary of the steps involved in a varying load and temperature test is given in Table 6. 34 Table 6 Representative procedure for a VLT test. Step Description 1 Place sample in the compression chamber 2 Tightly seal the chamber (cover rubber lining with grease) 3 Heat chamber to 500˚C at 100˚C /min 4 Perform three to four purges with Argon 5 Heat chamber up to 1300˚C at 100˚C /min 6 When temperature is stable, manually add weights to reach the first load 7 When the plot from register shows a steady slope, a load jump can be done 8 Continue to add/remove weights to reach the subsequent loads 9 For the temperature jump, heat chamber up to 1330˚C at 100˚C /min 10 When temperature is stable, manually adjust weights for load jumps 11 To perform the next load jump, lower the temperature in chamber back to 1300˚C at 100˚C /min 12 When temperature is stable, perform load jumps if needed 13 When the test is complete, lower the temperature in the chamber down to 20˚C 14 Allow the final temperature to be reached before removing sample from the chamber The measurements from the two sets of tests provided sufficient information necessary to calculate n and Q of the material during the creep test. Compressive creep testing was performed in the temperature range of 1200-1300 °C. Although this temperature range is well above the typical operating conditions of the SCWR, these temperatures relate to the “upset” conditions of the reactor, when it may be temporarily operated at elevated temperature. 3.6.4 Derivation of n and Q The raw experimental data were used in calculation of the creep parameters n and Q as follows: 3.6.4.1 Calculation of n An expression for strain rate and stress exponent can be obtained if two sets of experiments at a constant temperature are carried out. Assuming that the microstructure does not evolve considerably during the experiment and that the activation energy is constant, the equations for the strain rate during the two independent tests can be written as Equation 3a and 3b [20]: 35 ̇ ⁄ Equation 3a ̇ ⁄ Equation 3b The quotient of Equations 3a and 3b yields: ̇ ̇ ⁄ ⁄ = ⁄ ⁄ = Equation 4 Equation 4 allows to calculate the stress exponent between two load jumps from the measured strain rate during a test. 3.6.4.2 Calculation of Q Equation 1 was used to derive an expression for the activation energy which was evaluated from the data obtained during the constant load and varying temperature tests. In the work done by de Bernardi-Martín [20] it was shown that equation 1 can be rearranged to give: ( ( ̇ ) ( ) ) ( ) Equation 5 When Equation 5 is differentiated and a steady state deformation with constant load and stable microstructure is assumed, the following result is obtained [20]: ( ( ̇) ( ) ) ( ( ̇) ) Equation 6 Thus, during the temperature jump, Equation 6 takes the form [20]: ( ( ̇ ) ( ̇ ) ) Equation 7 The activation energy described by Equation 7 is expressed in units of J; however, Q is typically expressed in terms of J mol -1 . Thus, multiplying Boltzmann’s constant by 36 Avogadro’s number, correct units may be obtained and Equation 7 may be rewritten as follows (Equation 8) [20]: ( ( ̇ ) ( ̇ ) ) Equation 8 Where R = 8.315 J mol -1 . In order to calculate either stress exponent or activation energy the values for the strain rate and applied stress must be known. The strain rate was a function of the initial length, Lo, instantaneous length, Li, time, t, and the displacement of pistons in the machine, Y. Equation 5 for the strain rate at a given point in time is: ̇ ( ) Equation 9 The quotient of displacement and time is the slope from the register’s plot of displacement versus time. The change in length, ∆L, was obtained from the LVDT. The initial length of the sample was measured prior to testing. For the calculation of applied stress or instantaneous stress, σi, the classical expression was used: Equation 10 where F was the force, obtained from multiplication of the weights and the gravitational constant 9.81kg/ms 2 , and Si was the instantaneous cross section, which was estimated from the expression: Equation 11 37 The initial cross-sectional area, So, was measured prior to testing. The instantaneous strain, εi, was calculated from the initial and instantaneous lengths as shown: ( ) Equation 12 Once the necessary values were calculated, the stress exponent and activation energy can be obtained from equations 4 and 8. The load jump allowed the calculation of the stress exponent, while the temperature jump allowed for the calculation of the activation energy. For a detailed example of data calculation the reader is referred to Appendix A. 38 Chapter 4: Results and Discussion This section presents the results obtained during this research. For each composite, the densification behavior, general microstructure, chemical analysis and compressive creep results are discussed. 4.1 YSZ 4.1.1 Densification The as-sintered YSZ discs (Figure 12) fractured upon ejection from the sintering die. The surface of YSZ composites showed black coloration, which has been attributed to the diffusion of carbon from the graphite die and punches into the YSZ sample during SPS processing. Graphite contamination has been observed in SPS processing of other ceramic materials as well [71], [72]. Gentle surface grinding of the sample with emery paper was sufficient to remove the surface layer of graphite. The effect of SPS process parameters on the grain size and relative density of as-sintered composites is provided in Table 7. Table 7 Effect of SPS parameters on the grain size and relative density of YSZ composite. Composite Sintering Temperature (°C) Sintering Pressure (MPa) Sintering Time (min) Grain Size a (µm) Relative Density (%) YSZ 1300 30 7 2.6 ± 0.2 95.7 1400 30 7 10.0 ± 1.2 96.2 1500 50 10 24.0 ± 1.2 97.4 a -error shown taken from a 95% confidence interval As seen from Table 7, the sintering process parameters—particularly temperature—had a strong influence on the grain size of the composites. The largest grain size (~24 m) was recorded when the sintering temperature, pressure and time were highest. Further, there was a considerable grain growth in the sintered material compared to the starting particle size of the precursor powders (0.5 m) even at the lowest sintering temperature investigated. The matrix of the composites is shown in Figure 20, where the grain size growth is clearly visible. 39 Figure 20 General microstructure of YSZ composite processed at: a) 1300 °C, 30 MPa, 7 min, b) 1400 °C, 30 MPa, 7 min, c) 1500 °C, 50 MPa, 10 min. From Table 7 it is also observed that the density of the composites had a general tendency to increase with increasing sintering temperature. This indicates that dense composites can be manufactured using the SPS technique and achieve relative densities of >95 %. 40 4.1.2 Chemical Analysis and Microstructure The average composition of the matrix in the YSZ composite is provided in Table 8. Table 8 Chemical composition of matrix in YSZ composite. The XEDS results suggest that the average amount of Zr, Y and O was lower than the predicted theoretical amounts ( 63.8 wt% Zr, 10.8 wt% Y and 25.3 wt% O). The lower contents of Zr, Y and O were possibly due to the presence of impurities in the as-sintered material. For example, traces of Cl and Hf likely originated from the precursor powders (Table 1) and thus altered the relative weight fractions of the composites. The data in Table 8 also confirms that carbon diffused from the graphite die into the matrix during SPS. 4.1.3 XRD Analysis XRD scans for the YSZ composites are shown in Figure 21. Processing Parameters (°C/MPa/min) Wt% Zr Y O C Cl Hf 1300/30/7 61 8.4 21 8.2 <0.1 1.4 1400/30/7 60 8.3 19 10 <0.1 1.5 1500/50/10 62 8.5 21 6.1 <0.1 1.4 41 Figure 21 XRD plot for YSZ processed at: a) 1300 °C, 30 MPa, 7 min, b) 1400 °C, 30 MPa, 7 min, c) 1500 °C, 50 MPa, 10 min. The XRD results confirmed the retention of the cubic phase in all of the YSZ composites sintered with different SPS process parameters. Thus the desirable cubic phase was retained in YSZ with a sintering temperature as low as 1300 °C and a sintering time as short as 7 min. 4.1.4 Discussion The results of grain size measurements indicate that grain growth took place due to an increase of sintering temperature, as expected. By increasing the sintering temperature by 100 °C, the grain size grew by a factor of approximately five (~289 %). Also, the relative density increased by 0.5 % with an increase in sintering temperature. Bernard-Granger and Guizard’s [67] study on SPS of tetragonal YSZ (3YSZ) found a similar trend of a grain size increase of ~62 % and a relative density increase of ~5 %. According to Bernard-Granger and Guizard [67], the different magnitudes of growth between grain size and relative density (as a function of temperature) are common when the 42 relative density of SPS materials exceeds 95 %. Materials with high relative density contain small voids between grains, which are difficult to eliminate by compaction or grain boundary generation. The powder particles want to reduce the total surface energy through densification, but if the applied pressure is not sufficient at mobilizing the voids then formation of new grain boundaries is less likely than grain growth [67]. The YSZ materials in the present research sintered at 1300 and 1400 °C had relative densities of 98 % and 99 %, respectively. Therefore it is believed that the mechanism dominating densification in the present research experiments was mainly associated with grain growth. 4.2 YSZ doped with CeO2 4.2.1 Densification As-sintered CeO2-YSZ composites (Figure 13) had a tendency to fracture upon ejection from the sintering die. In relative terms, however, the CeO2-YSZ composites fractured more than the YSZ composite. The CeO2 –YSZ composites exhibited dark and orange colorations, as seen in Figure 13. Several studies [58], [71–73] have also reported similar orange coloration in CeO2 composites. The color variation was related to Ce 4+ reduction to Ce 3+ , enabled by carbon pickup due to the low oxygen partial pressure during sintering. This reduction induces a significant thermal lattice strain in the YSZ crystal lattice due to the ionic radius difference (Ce 4+ = 0.97 Å, Ce 3+ = 1.11Å and Zr 4+ = 0.84Å) and likely contributed to the cracking of the composites. Subsequent annealing of the as-sintered composites in a muffle furnace at 1000˚C for 1 hr resulted in a uniform pale white-yellow coloration, indicating that carbon impurities had decarburized and the Ce 3+ ions were oxidized to the Ce 4+ state. Table 9 shows the effect of CeO2 doping and SPS parameters on the relative density and grain size of the as-sintered composites. 43 Table 9 Effect of SPS parameters and doping level on the grain size and relative density of CeO2-YSZ composites. CeO2 (wt%) Temperature (°C) Pressure (MPa) Time (min) Grain Size a (µm) Relative Density (%) 6.4 1300 30 5 1.5 ± 0.2 95.7 12.7 1300 30 5 1.3 ± 0.1 89.1 18.8 1300 30 5 1.6 ± 0.1 88.6 1400 30 7 1.5 ± 0.2 97.3 a – error shown taken from a 95% confidence interval As seen from Table 9 the relative density of the composites ranged from 88 % to 97 %. By increasing the doping level of CeO2 from 6.4 wt% to 18.8 wt%, the relative density reduced, thereby suggesting that an increase in CeO2 addition hindered sintering of the composite. However, increasing the sintering temperature from 1300 °C to 1400 °C increased the relative density from 88 % to 97 %. Thus, the densification of CeO2-YSZ composites showed a stronger dependency on sintering temperature rather than on doping level. The grain size measurements provided in Table 9 suggest that increasing the doping level of CeO2 did not have a statistically significant effect on grain size. A representative micrograph of the annealed microstructure of CeO2-YSZ composites is shown in Figure 22. Figure 22 Representative SEM micrograph of 6.4 wt% CeO2 – YSZ after heat treatment. 44 4.2.2 Chemical Analysis and Microstructure 4.2.2.1 6.4 wt% CeO2-YSZ composite sintered at 1300 °C The microstructure of the 6.4 wt% CeO2-YSZ composite is shown in Figure 23, with XEDS results for the different regions included in Table 10. Figure 23 SEM micrograph of 6.4 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min: a) general microstructure, b) matrix, c) Ce-rich particle. 45 Table 10 Chemical composition of features in 6.4 wt% CeO2-YSZ composite sintered at 1300 °C, 30 Mpa, 5 min. Location Region of Interest Wt% Zr Y Ce O C Cl Hf A General Matrix 55 8.2 4.1 19 12 0.3 1.0 B Small Particle 23 1.9 42 9.2 13 9.9 0.4 C Large Particle Centre 0.2 0.2 67 7.8 9.2 15 <0.1 D Large Particle Interface 24 3.6 36 13 14 8.4 0.9 As seen from Table 10, the matrix of the composite contained a lower Ce content than the theoretical content of 6.4 wt%. This has been attributed to the presence of Cerium-rich regions found throughout the microstructure. As can be seen from the micrographs (Figure 23b and c), the CeO2-YSZ composite had localized Cerium-rich regions, classified by relative size as “small” and “large” particles. Small particles were observed along the grain boundaries and at the grain boundary triple points, while large particles were found randomly throughout the matrix. The large Cerium-rich particles consisted of two distinct areas as illustrated in Figure 24. The terminology presented in Figure 24 to describe the large Cerium- rich particles will be used throughout this chapter. Figure 24 Schematic of Cerium-rich particles in CeO2-YSZ composite. 46 The large particles (~19 μm in width and ~167 μm length) found in the 6.4 wt% CeO2- YSZ composite had a high concentration of Ce at the center region (“C” in Figure 23c), as well as a low Zr and Y content at the interface (“D” in Figure 23), compared to the general matrix. The data in Table 10 shows that the composition of the large particles was not uniform, with the cerium content decreasing from 67 wt% Ce at the center to 4 wt% at the particle-matrix interface. The small Cerium-rich particles, located between YSZ grains (Region “B” in Figure 23b), had an average Ce content of 4wt%, which while it was not as high as the Ce content at the center of large particles (67 wt%) it was still high compared to the Ce content in the general matrix (4.1 wt%). Together, both types of Cerium-rich particles resulted in a matrix that contained less Cerium than expected. It is believed that the Ce-rich particles (large and small) originated from insufficient blending of the precursor powders. In effect, the Ce-rich particles were possibly large agglomerates of CeO2 precursor powder. 4.2.2.2 12.7 wt% CeO2-YSZ composite sintered at 1300 °C The general microstructure of the as-sintered 12.7 wt% CeO2 –YSZ composite is shown in Figure 25, with XEDS results provided in Table 11. 47 Figure 25 SEM micrograph of 12.7 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min: a) general microstructure, b) matrix, c) Ce-rich particles, d) microstructure after heat treatment. Table 11 Chemical composition of features in 12.7 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min. Location Region of Interest Wt% Zr Y Ce O C Cl Hf A General Matrix 54 8.2 7.9 13 7.6 <0.1 1.2 B Small Particle 21 3.4 41 12 11 9.8 0.5 C Large Particle Centre 0 0 70 11 7.7 10 <0.1 D Large Particle Interface 40 5.6 26 18 7.9 0.6 0.9 48 The matrix of the 12.7 wt% CeO2 – YSZ composite had a lower Cerium content than the theoretical 12.7 wt%. Once again, presence of small and large Ce-rich particles was frequently observed, as shown by Figure 25b and c, and was likely the reason of the lower Cerium content in the matrix. The large particles were ~ 36 μm in width and ~186 μm in length. Large cavities (indicated by arrows in Figure 25c) were observed in the vicinity of Ce-rich particles, suggesting that these particles hindered sintering. The presence of these cavities also suggests that the SPS process parameters for a composite with this CeO2 concentration should be adjusted to enable enhanced densification. 4.2.2.3 18.8 wt% CeO2-YSZ composite sintered at 1300 °C The general microstructure of the as-sintered 18.8 wt% CeO2-YSZ composite is shown in Figure 26, with XEDS results provided in Table 12. 49 Figure 26 SEM micrograph of 18.8 wt% CeO2 – YSZ composite sintered at 1300 °C, 30MPa, 5 min: a) general microstructure, b) matrix, c) Ce-rich particles, d) microstructure after heat treatment. Table 12 Chemical composition of features in 18.8 wt% CeO2 – YSZ composite sintered at 1300 °C, 30 MPa, 5 min. Location Region of Interest Wt% Zr Y Ce O C Cl Hf A General Matrix 50 7.3 12 19 10 <0.1 1.2 B Small Particle 14 2.4 50 10 12 11 0.3 C Large Particle Centre 17 2.8 53 15 7.7 3.8 0.4 D Large Particle Interface 44 6.3 22 18 8 0.1 0.9 50 The matrix in the 18.8 wt% CeO2-YSZ composite contained Cerium-rich particles (Figure 26b,c), as observed in the previous composites. Large Cerium-rich particles had a size of ~23 μm in width and ~268 μm in length. Similarly as in the previous composite cavities at the extremities of the Ce-rich particles suggest that these agglomerares locally hindered sintering. The XEDS results shown in Table 12 show that the composition of the large particles was not uniform, as the center of the particle had higher Ce content (53 wt%) than the interface (22 wt%). Further, increasing the theoretical doping content from 12.7 wt% to 18.8 wt% increased the Ce content in the matrix from 7.9 wt% to 12 wt%. 4.2.2.4 18.8 wt% CeO2-YSZ composite sintered at 1400 °C The general microstructure of 18.8 wt% CeO2-YSZ sintered at 1400 °C, 30 MPa and 7 min is shown in Figure 27. XEDS results for the matrix and its constituents are summarized in Table 13. 51 Figure 27 SEM micrograph of 18.8 wt% CeO2 – YSZ composite sintered at 1400 °C, 30 MPa, 7 min: a) general microstructure, b) matrix, c) Ce-rich particles. 52 Table 13 Chemical composition of features in 18.8 wt% CeO2 – YSZ composite sintered at 1400 °C, 30 MPa, 7 min. Location Region of Interest Wt% Zr Y Ce O C Cl Hf A General Matrix 52 7.6 11 18 9.8 0.2 0.9 B Small Particle 19 3.2 44 11 11 10 0.3 C Large Particle Centre 0 <0.1 71 8.2 7.8 12 <0.1 D Large Particle Interface 28 3.8 44 14 8.3 0.9 0.7 The matrix in the 18.8 wt% CeO2 – YSZ composite had a lower Cerium content than the theoretical 18.8 wt%. Small and large Cerium-rich particles (Figure 27b,c) were found throughout the matrix, causing lower Cerium content in the matrix. The composition of the large Cerium-rich particles (~56 μm in width and ~350 μm in length) was not uniform, with the higher content of Ce in the center and lower at the interface. The Zr content in the center of the Ce-rich particles was negligible while at the interface it was measured at 28 wt%. 4.2.3 XRD The XRD scans for the CeO2-YSZ composites are shown in Figure 28. 53 Figure 28 XRD plot for CeO2 –YSZ composites sintered at 1300°C, 30 MPa, 5 min: a) 6.4wt% CeO2, b) 12.7wt% CeO2, c) 18.8wt% CeO2, d) 18.8wt% CeO2 – YSZ composite sintered at 1400°C, 30 MPa, 7 min. Performing peak analysis in XPert Pro software confirmed that the cubic phase of ZrO2 was present in all of the CeO2-YSZ composites. Therefore, addition of CeO2 did not de- stabilize the desirable cubic Zirconia matrix. 4.2.4 Discussion In all of the composites investigated, the matrix had lower cerium content than the corresponding theoretical amounts. This has been attributed to the formation of Cerium-rich particles within the composites. As the amount of CeO2 added to YSZ increased, the tendency to form Ce-rich particles increased. The change in grain size of the different CeO2 – YSZ composites was not directly related to the amount of CeO2 added. The grain size of the 6.4 wt% CeO2 – YSZ composite was ~13 % larger than in the case of 12.7 wt% CeO2-YSZ composite. However, further addition to the 18.8 wt% CeO2 level caused only ~14 % increase in grain size in comparison to the 12.7 wt% CeO2-YSZ composite. This trend is different to that observed by Boutz et al. [48], who found that the increase of CeO2 content between ~3 to 10 wt% did not change the grain size of 4 mol% YSZ sintered in the range of 1200-1400 °C. Thus, in the present research, the 54 addition of CeO2 in conjunction with evolution of thermal gradients in the composite during sintering likely affected the grain size. The highest relative density (~95 %) was obtained in the composite containing 6.4 wt% CeO2 sintered at 1300 °C. The density did not vary significantly with varying CeO2 concentration, but was affected by the sintering temperature, as expected. Large cerium-rich particles found in the CeO2 – YSZ composites had an elliptical geometry. The particle size of the CeO2 precursor powder was approximately 50-80 nm, while the size of the Ce-rich particles was ~950 nm. These observations suggest that these large particles could be groups of multiple (>10) powder particle agglomerates sintered together. SEM-XEDS line-scans of the large particles (Figure 29, Appendix B) were done to observe the possibility of diffusion of Cerium from the large particles into the matrix. Figure 29 Line-scan analysis of Cerium-rich particle in 18.8 wt% CeO2-YSZ composite. 55 As seen from Figure 29, the counts of Cerium on the sides of the particle were not identical, which is indicative that directional diffusion occurred. 4.3 YSZ doped with La2O3 4.3.1 Densification The as-sintered La2O3-YSZ composites (Figure 14) fractured, regardless of the doping level. However, the fracturing of La2O3-YSZ composites was significantly less extensive than in the case of the CeO2-YSZ composites. Therefore, the sinterability of the La2O3-YSZ composites was relatively easier than that of the CeO2-YSZ composites. The surface of the La2O3-YSZ composites had the same dark coloration from carbon pickup as the YSZ specimens. The isovalency of La 3+ with Y 3+ as well as the similarity in ionic sizes (Y 3+ =1.019Å, and La 3+ =1.172 Å) would favour the formation of a substitutional solid solution where the La 3+ ions are able to take the same position as the Y 3+ ions in the lattice. While there will still be an enlargement of the lattice due to this substitution of ions, the lattice strains may be of lower magnitude than in the case of the CeO2-YSZ composite, which could be possibly responsible for the reduced susceptibility to fragmentation. Table 14 shows the effect of La2O3 concentration and SPS parameters on the relative density and grain size of the as-sintered composites. 56 Table 14 Effect of SPS parameters and doping level on the grain size and relative density of La2O3 – YSZ composite. La2O3 (wt%) Temperature (°C) Pressure (MPa) Time (min) Grain Size a (µm) Relative Density (%) 11.6 1400 30 3 1.8 ± 0.1 ¥ 95.3 0.4 ± 0.1 § 1500 50 10 9.8 ± 1.2 ¥ 90.7 0.9 ± 0.2 § 21.6 1500 50 10 6.7± 0.5 ¥ 96.1 1.2 ± 0.2 § 30.5 1500 50 10 12.6 ± 1.4 ¥ 92.6 5.4 ± 0.6 ß 1.3 ± 0.2 § a -error shown taken from a 95% confidence interval ¥ -grain size corresponds to large grains in microstructure § -grain size corresponds to small grains in microstructure ß -grain size corresponds to large grains in secondary matrix The measured relative density for the La2O3-YSZ composite sintered at 1400 °C, 30 MPa and 3 min was 95 %. The measured relative density ranged from 90 % to 96 % for the La2O3-YSZ composites sintered at 1500 °C, 50 MPa and 10 min. It was also observed that the increase in doping level did not have a clear effect on the relative density at the 1500 °C sintering temperature. Figure 30 reveals that the microstructure had a bimodal character, consisting of two different grain families: i) large grains with equiaxed shape (marked by bold arrows) and ii) smaller grains located along the grain boundaries and triple points (marked by dashed arrows in Figure 30a). 57 Figure 30 Representative bimodal microstructure observed in La2O3 – YSZ composites: a) large and small grains indicated by arrows, b) small grain networks (dashed circles) containing voids (double arrows). The smaller grains were generally found along the grain boundaries and triple points of the larger grains, as seen in Figure 30a. When located at grain triple points, the small grains often formed circular networks that contained small voids in the center of the network, as seen in Figure 30b (indicated by double arrows). The data in Table 14 shows that by increasing the sintering temperature of the composites, the size of the large grains grew from ~1.8 μm to ~9.87 μm. Increasing the content of Lanthana in the composites from 11.6 wt% to 21.6 wt% had a grain refinement effect. However, upon addition of 30.5 wt% Lanthana to YSZ, the bimodal microstructure changed and had two large grain families and one small grain family. This unique microstructure will be discussed in detail in the next sections. 4.3.2 Chemical Analysis and Microstructure 4.3.2.1 11.6 wt% La2O3-YSZ composite sintered at 1400 °C The general microstructure of the as-sintered 11.6 wt% La2O3-YSZ composite sintered at 1400 °C is shown in Figure 31, with XEDS results provided in Table 15. 58 Figure 31 Microstructure of 11.6 wt% La2O3 – YSZ composite sintered at 1400 °C, 30 MPa, 3 min: a) as- sintered matrix, b) large grains, c) small grains, d) small grains at 100,000x magnification. 59 Table 15 Chemical composition of features in 11.6 wt% La2O3 – YSZ composite sintered at 1400 °C, 30 MPa, 3 min. Location Region of Interest Wt% Zr Y La O C Cl Hf A General Matrix 55 8.5 5.9 19 9.6 <0.1 1.1 - Large Grains 56 8.8 6 19 8.8 <0.1 1.3 - Small Grains 47 6.6 17 19 8.4 <0.1 1 - Small Grain Networks 51 7.6 11 19 9.3 - 1.1 As seen in the SEM micrograph in back-scattered electron mode (sensitive to Z-contrast) in Figure 31a, the small and large grains had a different color. From Table 15, it can be seen that the chemical composition of the larger grains was almost identical to that of the matrix (both having a grey color). The average composition of the small grains (light color), however, was 17 wt% La and 47 wt% Zr, while the large grains had a composition of 6 wt% La and 56 wt% Zr. In general, the small grains were found along the grain boundaries and grain triple points of the larger grains, but as it can be seen in Figure 31c and d, the small grains were also seen as part of larger networks throughout the general matrix. The chemical composition of the small grains along the grain boundaries (Figure 31b) was comparable to the chemical composition of the small grains that formed as large networks, as can be seen from Table 15. 4.3.2.2 11.6 wt% La2O3-YSZ composite sintered at 1500 °C The general microstructure of the as-sintered 11.6 wt% La2O3-YSZ composite sintered at 1500 °C is shown in Figure 32, with XEDS results provided in Table 16. 60 Figure 32 Microstructure of 11.6 wt% La2O3– YSZ composite sintered at 1500 °C, 50 MPa, 10 min: a) general microstructure, b) Lanthanum-rich particles, c) large grains, d) small grains. 61 Table 16 Chemical composition of features in 11.6 wt% La2O3 – YSZ composite sintered at 1500 °C, 50 MPa, 10 min. Location Region of Interest Wt% Zr Y La O C Cl Hf A General Matrix 64 9.7 3.6 16 4.8 <0.1 1.5 B Particle Interface 30 3.8 51 11 2.6 0 0.7 C Particle Center 0.15 0.14 75 5.4 3.0 15 <0.1 - Large Grains 58 8.8 3.6 18 9.4 <0.1 1.4 - Small Grains 48 6.4 16 18 8.9 <0.1 1.1 - Small Grain Networks 48 6.8 14 19 9.2 <0.1 1 This composite (Figure 32a) had a bimodal microstructure (Figure 32c) composed of large and small grains. As in the previous composite, the small grains were commonly found along the grain boundaries and grain triple points of the larger grains, as shown in Figure 32c and d. Although the matrix of this composite (Figure 32a) had a relatively homogeneous chemical composition, large Lanthanum-rich particles with elliptical geometry were observed throughout the matrix, as shown in Figure 32 b. The size of these particles ranged from ~100 to 250 µm in length, and a width up to ~125 µm. Line scans done using SEM-XEDS revealed that the concentration of Lanthanum, Zirconium, and Yttrium across the particles was not constant. The center of these particles (Region “C” in Figure 32b) had higher La and lower Zr and Y contents than the interface. This further supports the possibility that this elliptical particles were the La2O3 powder particles , which after being pressed during sintering diffused out into the matrix due to the high temperatures used during SPS processing. The particle size of the raw La2O3 powder was ~80 nm, but the La-rich agglomerates were as large as ~10 µm (Figure 11c). 4.3.2.3 21.6 wt% La2O3-YSZ composite sintered at 1500 °C The general microstructure of the as-sintered 21.6 wt% La2O3-YSZ composite sintered at 1500 °C is shown in Figure 33, with XEDS results provided in Table 17. 62 Figure 33 Microstructure of 21.2 wt% La2O3 – YSZ composite sintered at 1500 °C, 50 MPa, 10 min: a) general microstructure, b) Lanthanum-rich particles, c) large grains, d) small grains. 63 Table 17 Chemical composition of features in 21.6 wt% La2O3 – YSZ composite sintered at 1500 °C, 50 MPa, 10 min. Location Region of Interest Wt% Zr Y La O C Cl Hf A General Matrix 64 9.2 6.3 14 3.6 <0.1 1.6 B Particle Interface 28 3.4 45 8.3 1.6 0.1 0.7 C Particle Center 4.9 0.83 72 5.8 2.8 13 0.2 - Large Grains 48 7.5 3.2 22 18 - 0.9 - Small Grains 47 6.7 7.2 23 14 <0.1 0.8 - Small Grain Networks 40 6 12 23 18 - 0.8 As in the previous composite, the 30.5 wt% La2O3-YSZ the larger grains (Figure 33c) had a composition comparable to the matrix, while the smaller grains (Figure 33d) had higher La and lower Zr and Y contents than the matrix. Line scans done on these small grains (Figure 34) confirmed the higher La and lower Zr and Y contents in the small grains whether the small grains were located in a triple point or between two larger grains. From Figure 34, it can also be seen that the small grains appear to be pining adjacent larger grains. Areas without small grains usually contained a void. 64 Figure 34 SEM-XEDS line-scan analysis of small grains in 21.6 wt% La2O3-YSZ. The large La-rich agglomerates seen in the 11.6 wt% La2O3-YSZ composite were observed again in the 21.6 wt% La2O3-YSZ composite. The shapes of these particles were variants of ellipses, and many of the particles seemed to connect with neighboring particles as shown in Figure 33b. Line-scans (Figure 35) revealed that the chemical composition of such particles was not constant, with the largest variation being the concentration of 65 Lanthanum. The XEDS analysis of the different regions within the particles (Table 17) shows that the center of the particles had an average of 72 wt% La, which decreased to ~45 wt% La in the interface region and then further decreased to 6.3 wt% La in the matrix. Figure 35 SEM-XEDS line-scan analysis of large particle in 21.6 wt% La2O3-YSZ. 4.3.2.4 30.5 wt% La2O3-YSZ composite sintered at 1500 °C The general microstructure of the as-sintered 30.5 wt% La2O3-YSZ composite sintered at 1500°C is shown in Figure 36, with XEDS results provided in Table 18. 66 Figure 36 Microstructure of 30.5 wt% La2O3 – YSZ composite sintered at 1500 °C, 50 MPa, 10 min: a) microstructure with primary (“A”) and secondary (“B”) matrices, b) large grains in primary matrix, c) large grains in secondary matrix, d) small grains in primary matrix. 67 Table 18 Chemical composition of features in 30.5 wt% La2O3 – YSZ composite sintered at 1500 °C, 50 MPa, 10 min. Unlike the previous composites, this material’s microstructure showed an additional family of grains, as shown by Figure 36a. The 30.5 wt% La2O3-YSZ composite had two groups of large grains and one group of small grains. The average grain size of the primary matrix grains (Figure 37a) was ~12.6 μm ± 0.16 m, while the grain size of the secondary matrix grains (Figure 37b) was ~5.3 μm ± 0.61 m. The small grains found in both areas (Figure 38) had an average grain size of ~1.3 μm ± 0.20 m. Therefore, based on the XEDS results (Table 18), the microstructure can be thought of as having a primary matrix and a secondary matrix, both containing large and small grains. Location Region of Interest Wt% Zr Y La O C Cl Hf A General Matrix 63 9.3 10 11 3.9 <0.1 1.5 B Secondary Matrix 34 4.4 45 11 3.5 <0.1 1 - Large Grains 1 (grey) 49 7.6 5.6 25 11 <0.1 0.9 - Large Grains 2 (white) 29 3.8 34 22 9.2 <0.1 0.7 - Small Grains 37 4.9 21 24 11 <0.1 1 - Small Grain Networks 32 5 12 26 23 - 0.6 68 Figure 37 Large grains (bold arrow) observed in: a) primary matrix, b) secondary matrix of 30.5 wt% La2O3 – YSZ composite. Figure 38 Small grains (dashed arrows) observed in: a) primary matrix, b) secondary matrix of 30.5 wt% La2O3 – YSZ composite. 69 The secondary matrix contained more La than the primary matrix. The size of the large grains in the secondary matrix (5.36 μm) was less than half the size of the large grains in the primary matrix (12.6 μm). The differences in grain size and composition suggests that La2O3 diffused more effectively into some areas, as indicated by the line scan in Figure 39, which later established the secondary matrix. Figure 39 SEM-XEDS line-scan analysis of secondary matrix in 30.5 wt% La2O3-YSZ composite showing lower Zr and higher La counts than in primary matrix. 4.3.3 XRD Analysis The XRD scans for La2O3-YSZ composites are shown in Figure 40. 70 Figure 40 XRD plot for: a) 11.6 wt% La2O3 –YSZ sintered at 1400 °C, 30 MPa, 3 min, b) 11.6 wt% La2O3 –YSZ sintered at 150 0°C, 50 MPa, 10 min, c) 21.6 wt% La2O3 –YSZ sintered at 1500 °C, 50 MPa, 10 min, d) 30.5 wt% La2O3-YSZ sintered at 1500 °C, 50 MPa, 10 min. All La2O3-YSZ composites retained the cubic phase of ZrO2. In addition to the cubic phases of Zirconia, some secondary phases were also detected in the Lanthana-containing composites. These secondary peaks could belong to the pyrochlore phase La2Zr2O7 or the Lanthanum Tri-hydroxide compound La(OH)3. Evolution of the pyrochlore La2Zr2O7 phase was previously reported for 33 mol% (~56 wt%) La2O3 content in ZrO2 processed at 1550 °C [74]. 4.3.4 Discussion The data presented in Table 14 shows the effect of La2O3 content on the grain size and relative density. In the case of relative density, it was not possible to establish a direct relationship between La2O3 content or increase in sintering temperature and the relative density. In all of the composites, however, the relative density exceeded 90 %. The grain size of the general matrix was reduced from 9.8 μm to 6.7 μm when the concentration increased from 11.6 wt% to 21.6 wt% La2O3, but further addition to 30.5 wt% La2O3 resulted in a microstructure with two distinct matrices, each containing different sized- 71 grains (bimodal microstructure). The large grains in the primary matrix were ~12.6 μm in width, while the large grains in the secondary matrix were ~5.3 μm. The small grains in the bimodal microstructure of all the La2O3-YSZ composites had a tendency to increase in size with addition of La2O3, increasing from 0.4 μm in the 11.6 wt% La2O3-YSZ composite to 1.3 μm (~40% increase) in the 30.5 wt% La2O3-YSZ composite. The overall grain refinement effect of La2O3 in YSZ has been reported in literature [51], [56]. A possible explanation for this effect was provided by Liu et al. (2009), who found that Lanthanum hindered diffusion of ZrO2 in La2O3-coated YSZ powders. Since La 3+ has a large ionic radius, this results in the formation of a solid solution with an enlarged lattice and hinders Zirconium diffusion and grain growth. Tsipas [43] discussed the competitive diffusion of Lanthanum and Yttrium when both of these elements are present in ZrO2 coatings, which could be responsible for the inhibition of grain growth in the 30.5 wt% La2O3-YSZ composite studied here. The presence of small grains and voids close to the small grains in the bimodal microstructure observed in the 11-30.5 wt% La2O3-YSZ composites was also observed in conventionally sintered La2O3-YSZ composites [56] and La2O3-coated YSZ [51–53], [58]. The changes in the microstructure reported in the literature and in the present work seem to be a strong function of processing parameters, such as sintering temperature and time. The microstructure of the La2O3-YSZ composites investigated in this work also contained large Lanthanum-rich particles or regions spread throughout the matrix. These particles could have been produced from the agglomeration and subsequent pressing of the La2O3 precursor powder during sintering of the composite material. The particle size of the La2O3 powder was approximately 80 nm, while the size in the particle regions was ~2 μm, which would correspond to a group of multiple (>20) powder particle agglomerates. In the 30.5 wt% La2O3-YSZ composite, the morphology of the Lanthanum-rich particles was unique; the Lanthanum-rich regions were larger and lacked a concentrated “particle center” or an elliptical shape. It is possible that the La2O3 powder agglomerates diffused more effectively into the general matrix, restricted primary grain growth, and thus appeared to form a secondary matrix in the microstructure. Line-scans of these particles confirmed that diffusion of Lanthanum into the general matrix occurred. The line-scans in Figure 35 and Figure 39 showed that the Lanthanum counts were higher at the center of the 72 particle and decreased towards the matrix area on each side. The counts of Lanthanum on each side of the particle were not identical, indicating that directional diffusion occurred. 4.4 YSZ doped with Nd2O3 4.4.1 Densification The as-sintered Nd2O3-YSZ composites were shown in Figure 15. The Nd2O3-YSZ composites had few fractures. The isovalency of Nd 3+ with Y 3+ and the similarity in ionic sizes (Y 3+ =1.019Å , Nd 3+ =1.123 Å) favour the formation of a substitutional solid solution. The enlargement of the lattice due to the substitution of ions was possibly enough to generate internal strains that promoted cracking. The Nd2O3-YSZ composites had dark surface due to carbon pickup as was seen in the case of other YSZ composites. Table 19 shows the effect of Nd2O3 concentration on the relative density and grain size of the composites. Table 19 Effect of doping level on the grain size and relative density of Nd2O3 – YSZ composite. Nd2O3 (wt%) Sintering Temperature (°C) Sintering Pressure (MPa) Sintering Time (min) Grain Size a (µm) Relative Density (%) 12 1400 40 5 2.7 ± 0.2 ¥ 97.3 0.2± 0.0§ 22.2 1400 40 5 2.4± 0.2 ¥ 96.5 0.2± 0.1§ a -error shown taken from a 95% confidence interval ¥ -grain size corresponds to large grains in microstructure § -grain size corresponds to small grains in microstructure The data provided in Table 19 suggests that increasing the Nd2O3 content in the composites reduced the relative density by <1 %. This suggests that addition of this oxide can be carried out without compromising the density of the composite. In terms of the grain size, the Nd2O3-YSZ composites exhibited a bimodal microstructure, which is illustrated in Figure 41. 73 Figure 41 Representative bimodal microstructure observed in Nd2O3 – YSZ composites: a) large grains shown by bold arrow, b) small grains shown by dashed arrow. The data in Table 19 suggest that the large grains in the microstructure experienced a size growth of ~10 % as a result of increased Nd2O3 content. In contrast, in the case of the small grains the increase of Nd2O3 content did not significantly affect the grain size. 4.4.2 Chemical Analysis and Microstructure 4.4.2.1 12 wt% Nd2O3-YSZ composite sintered at 1400 °C General microstructure of the 12 wt% Nd2O3-YSZ composite is shown in Figure 42, with XEDS results provided in Table 20. 74 Figure 42 SEM micrograph of 12 wt% Nd2O3 – YSZ composite sintered at 1400 °C, 40 MPa, 5 min: a) as- sintered matrix, b) bimodal microstructure, c) large grains, d) small grains 75 Table 20 Chemical composition of features in 12 wt% Nd2O3 – YSZ composite sintered at 1400 °C, 40 MPa, 5 min. Location Region of Interest Wt% Zr Y Nd O C Cl Hf A General Matrix 54 8.4 6.5 21 9 0.2 1.2 - Large Grains 53 7.8 6.3 21 10 <0.1 1.1 - Small Grains 43 6.7 15 23 11 <0.1 0.9 The as-sintered matrix, shown in Figure 42a, had a bimodal microstructure as previously discussed in Section 4.4.1. The bimodal microstructure was composed of large grains (“A” in Figure 42b) and small grains (“B” in Figure 42b). As seen in Table 20, the chemical composition of the large grains, shown in Figure 42c, was comparable to that of the matrix, while the small grains (Figure 42d) had higher Neodymium content (15 wt%), lower Zirconium (43 wt%) and Yttrium contents (6.7 wt%). The micrograph in Figure 42c and d show the formation of networks between the small grains, which were measured as small as 100 nm (Figure ) and were always accompanied by small voids in between. In contrast, the large grains (Figure 42b) had equiaxed geometry throughout the matrix. 4.4.2.2 22.2 wt% Nd2O3-YSZ composite sintered at 1400 °C General microstructure of the 22.2 wt% Nd2O3-YSZ composite is shown in Figure 43, with XEDS results provided in Table 21. 76 Figure 43 SEM micrograph of 22.2 wt% Nd2O3 – YSZ composite sintered at 1400 °C, 40 MPa, 5 min: a) as-sintered matrix, b) bimodal microstructure, c) large grains, d) small grains. Table 21 Chemical composition of features in 22 wt% Nd2O3 – YSZ composite sintered at 1400 °C, 40 MPa, 5 min. Location Region of Interest Wt% Zr Y Nd O C Cl Hf A General Matrix 51 8 12 19 9 0.2 1.1 - Large Grains 48 7.6 12 21 10 <0.1 1 - Small Grains 40 5.9 20 22 9.7 <0.1 0.8 The as-sintered matrix of this composite (Figure 43a) had a bimodal microstructure composed of large grains (“A” in Figure 43b) and small grains (“B” in Figure 43b). As shown by Table 21, the chemical composition of the large grains (Figure 43c) was 77 comparable to that of the matrix, while the small grains (Figure 43d) had higher Neodymium content (20 wt%), but lower Zirconium (40 wt%) and Yttrium contents (5.9 wt%). The network arrangement of the small grains in the 22.2 wt% Nd2O3-YSZ composite (Figure 43c) was the same as the one observed in the 12 wt% Nd2O3-YSZ composite. 4.4.3 XRD Analysis The XRD scans for the Nd2O3-YSZ composites are shown in Figure 44. Figure 44 XRD plot for: a) 12 wt% Nd2O3 –YSZ sintered at 1400 °C, 40 MPa, 5 min, b) 22.2 wt% Nd2O3 –YSZ sintered at 1400 °C, 40 MPa, 5 min. Both composites retained the cubic phase of ZrO2. The Nd2O3-YSZ composite scans show one additional peak for the 12 wt% Nd2O3-YSZ composite at a 2ϴ value of ~28°. Due to limitations of the XPert Pro software, this peak could not been identified and could be a result of impurities trapped in the precursor powder. In general, the XRD data shows that addition of up to 22 wt% Nd2O3 is possible without compromising the stability of cubic ZrO2. 78 4.4.4 Discussion The data presented in Table 19 show the effect of Nd2O3 on the grain size and relative density. By increasing the Nd2O3 concentration in the Nd2O3-YSZ composite from 12 to 22.2 wt%, the size of the large grains decreased by ~10 % with the addition of 22 wt% Neodymium, while the size of the small grains increased by ~20 %. The relative density of the material decreased minimally (<0.1 %). The observable changes in the microstructure are limited to a higher number of small grains and a higher concentration of Neodymium in both the matrix and the small grains. The bimodal nature of the composites investigated in this work has been reported in literature [14], [63], but the effect of varying Nd2O3 concentration on YSZ remains unclear. While Xu et al. (2004) did not report grain growth when increasing Nd2O3 content from 1.5 to 2 mol% in conventionally sintered 1.5 mol%Y2O3-ZrO2, Salehi et al. [63] did find grain size reduction when increasing Nd2O3 content from 1 to 2 mol% in 1 mol%Y2O3-ZrO2 processed through SPS. Salehi et al. also commented on competitive dissolution of Nd2O3 with Y2O3, suggesting that dissolution of Nd2O3 in ZrO2 takes more time and requires higher temperatures than that of Y2O3. It appears that different sintering times of the ZrO2-Y2O3-Nd2O3 materials (3 min in SPS), like those investigated by Salehi et al., versus conventional sintering (>60 min in conventional sintering), like those investigated by Xu et al., influenced the dissolution rates of each of the rare earth oxides and consequently grain growth. Although the composite was sintered for a longer period of time, the microstructure did not experience substantial grain growth. A possible explanation could be that during SPS, one of the two rare earth oxide’s dissolution was favored over the other during the short sintering time. The results in the research suggests that the composition of 12-22 wt% Nd2O3-YSZ is likely within two solubility lines and consequently the microstructure and phase stability remain fairly constant. The lack of a ZrO2-Y2O3-Nd2O3 phase diagram hinders further understanding of this composite. 4.5 Compressive Creep Results Compressive creep testing was carried out in two steps, as described in Chapter 3 in detail. First, a uniaxial compression test at constant cross-head speed was performed. Since 79 the plots obtained from this test were in the form of a paper trace (i.e., analog recording on a roll of paper), the results are not included in this thesis. The second step of compressive creep testing involved varying the load and temperature to enable calculation of the value of n and Q. the results of these experiments are discussed in the following sections. 4.5.1 YSZ Compressive creep testing was carried out with YSZ composite sintered at 1400 °C and 30 MPa for 7 min. This composite was selected for testing, as at these SPS process parameters enabled fabrication of a dense sample suitable for compression testing. The compressive creep was performed in the temperature range of 1300-1330 °C, as the maximum testing temperature had to remain at least 50 °C below the sintering temperature to avoid evolution of the microstructure. A starting load of 133 MPa was applied, as this stress was determined (form uniaxial compression test at constant cross-head speed) as the stress required to initiate plastic deformation in the composite. The strain rate vs. strain plot for the YSZ sample is shown in Figure 45. Figure 45 Compressive creep response of YSZ sintered at 1400 °C, 30 MPa, 7 min. As the plot shows, the test started at a temperature of 1300 °C and a stress of 134 MPa (~660 N). This condition was maintained until a steady state condition was achieved 80 (observed as a plateau on the graph). Then, the load increased to 153 MPa and the system was allowed to achieve a new steady state. By varying the stress and observing the strain rate response, a stress exponent n was calculated and is indicated above the vertical dotted line. The activation energy was calculated in a similar manner, but instead of load, the temperature was varied. The two temperature jumps are indicated in Figure 45 by the two bolded dotted vertical lines. For each temperature jump an activation energy Q value was calculated and is indicated above the bolded dotted vertical line. Each dotted vertical line in Figure 45 separates a regime with different applied stress. As can be seen from the results, the stress exponent ranged from 1.7 to 2.7. Typically, a stress exponent in the range of 1-2 is associated with lattice diffusion creep. At n=2 grain boundary sliding dominates and at n=3 dislocation climb becomes operative [12], [13], [16], [19], [22], [28]. Almost all of the measured n values were non-integers, suggesting that different mechanisms were active simultaneously or become dominant in a competitive process [18], [25], [75]. Thus, the measured values of n in this work suggest that the creep rate of YSZ was likely controlled by both lattice diffusion and grain boundary sliding simultaneously. The average activation energy of the YSZ composite was 730 kJ/mol in the temperature range between 1300-1330 °C. This is a relatively high value compared to typical values found in literature (~500 kJ/mol) [18], [19], [24], [28]. Usually the activation energy value would be compared to the activation energies of different creep mechanisms to determine the active mechanism. However, the value measured in this work was significantly higher than the activation energy of the individual creep mechanisms, which lie between 300-500 kJ/mol [13], [22], [36] . A possible explanation could be the cumulative nature of multiple mechanisms [36], such as diffusion of two different ions. In the case of YSZ, the diffusion of both Y 3+ and Zr 4+ ions could be taking place, which might explain the relatively high activation energy. 4.5.1.1 Microstructure of YSZ after creep The grain size of the YSZ composite before the creep test was ~10.0 μm. After the creep test, the average grain size decreased to 3.7 μm ± 0.2 μm, as can be observed from Figure 46. This grain size reduction was not expected, since high temperature exposure typically promotes grain growth. There is no published literature on creep in YSZ that resulted in grain 81 size reduction. Therefore, the present research suggests that stress-induced grain transformation, or recrystallization, may have occurred in this material. Therefore, high stress associated with the total strain of 16 % (corresponding to ~788 μm based on the initial length of 4.93 mm) may have contributed to grain transformation. Figure 46 SEM back-scattered micrographs of YSZ composite: a) before creep, b) after creep. Examination of the post-creep YSZ composite revealed small, round particles homogeneously distributed throughout the matrix as indicated by arrows in Figure 47b. 82 Figure 47 SEM secondary-electron micrographs of YSZ: a) before creep, b) after creep. These particles were not present in the original YSZ composite (Figure 47a). The average width of these particles was measured to be 0.8 μm ± 0.1 μm. SEM-XEDS analysis (Table 22) revealed that these particles had a lower Zr and Y contents (~6 % and 33 %, respectively), but a higher O content (~19 %), than the matrix. These trends were also observed in the results of line-scans, as seen in Figure 48. Table 22 Chemical composition of features in YSZ composite before and after creep. Region of Interest Wt% Zr Y O C Cl Hf Matrix (Prior to creep) 60 8.3 19 10 <0.1 1.5 Matrix (Post-creep) 62 8.4 17 10 <0.1 1.5 Particle (Post-creep) 58 6.3 21 12 <0.1 1.2 83 Figure 48 Line-scans of particles observed in post-creep YSZ microstructure: a) analyzed site with labeled areas of interest, b) line scan of particle 1, c) line scan of particle 2. The line-scan of particles (Figure 48) suggests that Zr and Y may have diffused from the particle towards the matrix, while the oxygen concentration was highest near the center of the particle. It is possible that the particle be composed of a metastable tetragonal phase (t’) with oxygen displaced along the c-axis of the crystal lattice. Nazarpour et al.[76] reported such phenomenon in YSZ doped with Al2O3, where oversaturation of YSZ with Al2O3 may cause a cubic to tetragonal transformation and hence the formation of the t’ phase. In the present work, performing XRD analysis would be necessary to verify whether the particles observed 84 in post-creep YSZ represent polymorphs of Zirconia such as the one discussed by Nazarpour et al. In addition to the formation of the particles, the micrographs in Figure 46 and Figure 47 revealed presence of voids along the grain boundaries and at grain boundary triple points after creep. The cracks observed in the YSZ composite investigated in this research had similar crack geometry, with straight edges clearly belonging to the grain boundaries through which the crack traveled as shown in Figure 49. Crenellated cracks have been frequently observed in creep studies of cubic Zirconia [18], [19], and were associated with grain boundary sliding. Figure 49 SEM micrographs of post-creep YSZ composite: a) crenellated cracked boundaries, b) trans- particle fracturing (arrows 1 to 3) and solid particle (arrow 4). It was observed that the primary crack propagated along the grain boundaries, suggesting that these regions were the weakest part of the microstructure. Further, it was seen that secondary cracks also fractured the particles (trans-particle fracture, as seen in particles 1,2, and 3 in Figure 49b) while in other cases the cracks propagated around particles on the grain boundaries (particle 4 in Figure 49b). If the particles were in fact tetragonal or monoclinic YSZ, then the deviations of crack path could be the result of a variation of 85 ductility and hardness in the matrix. Similarly, the grain boundary strength could be locally affected. The variation on the stress exponent and the high values for activation energy could in fact be reflecting changes associated with the microstructure inhomogeneity during phase changes, whether it is transformation toughening by cubic-to-tetragonal Zirconia transformation or by weakening of the boundaries by the presence of monoclinic Zirconia. As seen from Figure 49a, the post-creep microstructure also contained voids. These voids, with aan average width of 1.5 μm ± 0.12 μm, were not observed in the microstructure prior to creep. Had these voids diffused along the grain boundaries as in Cobble creep, the distance to travel would be relatively long. Instead, it appears that void formation was likely associated with grain boundary sliding [18], [22], since majority of the grains in the YSZ material after creep exposure show cracked boundaries. Some of the voids were likely caused by particles falling out of the matrix. This mechanism would be feasible if the particles were in fact nuclei of monoclinic Zr, which would mean that a volume expansion took place during creep testing, causing the observed bulging and cracking of the matrix in the vicinity of the particles, as seen in Figure 47b. 4.5.2 6.4 wt% CeO2-YSZ The varying load and temperature compressive creep test of 6.4 wt% CeO2-YSZ was completed in the temperature range of 1200-1230 °C and 78-249 MPa. The strain rate vs. strain plot obtained from the series of varying load and temperature tests is shown in Figure 50. 86 Figure 50 Compressive creep response of 6.4 wt% CeO2-YSZ processed at 1300 °C, 30 MPa, 5 min. The stress exponent ranged from 1.2 to 2.7, while the activation energy was 411 kJ/mol at 1230 °C and 572 kJ/mol at 1200 °C. The sample was deformed to a strain of ~16 %, which corresponds to 752 μm based on the initial length of 4.7 mm. 4.5.2.1 Microstructure of 6.4 wt% CeO2-YSZ after creep SEM micrographs of the composite’s matrix before and after creep are shown in Figure 51 and chemical analysis of the microstructure after creep is shown in Table 23. 87 Figure 51 SEM micrographs of general microstructure of 6.4 wt% CeO2-YSZ: a) before creep, b) after creep. Table 23 Chemical composition of features in 6.4 wt% CeO2-YSZ composite before and after creep. As seen from Table 23, the chemical composition of the matrix remained almost unchanged after compressive creep. However, the grain size of the material after compressive creep was 3.7 μm ± 0.2 μm, indicating a 147 % increase in size. In addition to this substantial grain size reduction, two unique features of the post-creep microstructure of interest were: i) dispersed Cerium-rich particles, and ii) cracked grain boundaries. The large (~19 μm in width and ~167 μm length) Cerium-rich particles that were present in the composite before creep testing (Figure 23c) were not present in the composite after creep at 1230 °C. However, smaller Cerium-rich particles, as shown by Figure 52, with an Region of Interest Wt% Zr Y Ce O C Cl Hf Matrix (Prior to Creep) 55 8.2 4.1 19 12 0.3 1 Matrix (Post Creep) 54 8.3 4.5 21 9.7 <0.1 1.2 Small Particle (Prior to Creep) 23 1.9 42 9.2 13 9.9 0.4 Dispersed Particle (Post Creep) 31 <0.1 45 12 4.9 <0.1 0.7 88 average size of 3.2 μm ± 0.2 μm were seen dispersed in random regions of the microstructure. Figure 52 SEM micrographs of Cerium-rich particles (arrows) in post creep 6.4 wt% CeO2-YSZ. The average chemical composition of the small Ce-rich particles shown in Figure 52 are also included in Table 23. These dispersed Ce-rich particles had a comparable chemical composition to the small Cerium-rich particles that were observed prior to creep (Figure 23b). Close examination of the surface of the Ce-rich particles revealed that some of the particles had an apparent surface texture, as can be seen in Figure 53. 89 Figure 53 Cerium-rich particles in post creep 6.4 wt% CeO2-YSZ: a) SEM micrograph, b) screenshot of SEM-XEDS analysis software showing close up of textured vs. non textured particles. No Visible Texture No Visible Texture Visible Texture Visible Texture 90 This “texture” was identified by presence of small dark areas on the surface of the white- coloured particles, as shown in Figure 53b. A closer look to these Ce-rich particles revealed a possible correlation between the observed apparent surface texture and particle-crack interaction. As seen in Figure 54, two distinct modes of interaction with a propagating crack were observed. Figure 54 Cerium-rich particles: a) trans-particle fracture, b) inter-particle fracture. A trans-particle fracture is shown in Figure 54a; as shown by this image, this fracturing was typically accompanied by some degree of detachment of the particle’s edges from the matrix. In other instances, however, a crack would simply be deflected from its path and forced to go around the particle (inter-particle fracture) as shown in Figure 54b. Detachment of the particles from the grain boundaries is shown in Figure 54a, a behaviour observed in nearly all particles. In addition to the different types of fractures experienced by the particles, their chemical composition was different depending on the apparent surface texture they presented showing apparent surface texture was different to particles without it. This information is summarized in Table 24. 91 Table 24 Chemical composition of Cerium-rich particles in post-creep 6.4 wt% CeO2-YSZ composite. From the chemical analysis results, it appears that there are some particles in the material with higher Cerium content that did not fracture. The high concentration of Cerium may result in enhanced mechanical properties, thereby making the particles a more difficult obstacle for a crack to cut through, as it is the case in coherent particle dispersion. There is a paucity of published data on compressive creep of CeO2-YSZ composites. The most relevant work to the present research (for purpose of comparison) was completed by de Bernardi-Martín [20]. In his study, 12 mol% CeO2-ZrO2 was tested in compressive creep at 1250 °C and 50 MPa. The calculated stress exponent was ~2, and the activation energy was 550 kJ/mol. Similar to the present research, de Bernardi-Martín [20] observed grain growth (~29 %) after creep. For the tests carried out in the study presented here, the stress exponent exceeded a value of 2, which suggests that grain boundary and possibly additional mechanisms were active to accommodate compressive creep. The micrograph in Figure 51b shows that the majority of the grain boundaries indeed cracked, and these cracks typically had a crenellated shape, which has been associated with high temperature grain boundary sliding [18], [19]. Also, the activation energy could have been affected by the cracking of Cerium-rich particles. It can be deduced that the co-stabilization of cubic Zirconia with Yttrium and Cerium oxides does not negatively affect the creep resistance of the material. The calculated n values, the crenellated cracking of grain boundaries and the cavitation at the grain triple points all suggest that grain boundary sliding was activated in temperature range of 1200-1250 °C and stress range of 50 to 249 MPa. To obtain a conclusive evidence for identification of the creep mechanism, as well as the possibility for transition or superposition of mechanisms [15], Wt% Zr Y Ce O C Cl Hf Visible texture 29 5.0 41 16 7.4 <0.1 0.6 No visible texture 19 3.5 56 13 6.3 <0.1 0.5 Trans-particle fracture 24 4.2 49 15 6.5 <0.1 0.6 Inter-particle fracture 19 3.7 54 16 5.9 <0.1 0.3 92 [22], [69], TEM and XRD analysis of secondary phases and impurities, ionic species diffusion, and particle dispersion would be necessary. 4.5.3 11.6 wt% La2O3-YSZ To the best of the author’s knowledge, there is no available literature on the behaviour of compressive creep of La2O3-YSZ composites. Thus, the following results are of pioneering nature and future work is necessary to be carried out for further analysis of the results. The varying load and temperature compressive creep test of YSZ was completed in the temperature range of 1300-1330 °C and 45-78 MPa. The results are plotted in Figure 55. Figure 55 Compressive creep response of 11.6 wt% La2O3-YSZ processed at 1400 °C, 30 MPa, 3 min. From Figure 55 it can be seen that the calculated stress exponent range was 1.7-2.1, and the activation energy was 546 kJ/mol at 1330°C and 586 kJ/mol at 1300 °C. The sample was deformed to a strain of ~12 %, which corresponds to approximately 551 μm based on the initial length of 4.59 mm. 93 4.5.3.1 Microstructure of 11.6 wt% La2O3-YSZ after creep SEM micrographs of the bimodal microstructure before and after creep testing are provided in Figure 56, with corresponding XEDS results provided in Table 25. Figure 56 SEM micrographs of 11.6 wt% La2O3-YSZ: a) before creep, b) after creep. Table 25 Chemical composition of features in 11.6 wt% La2O3 – YSZ composite before and after creep. Region of Interest Wt% Zr Y La O C Cl Hf Large Grains (Prior to Creep) 56 8.8 6 19 8.8 <0.1 1.3 Large Grains (Post Creep) 57 8.7 6.2 15 11 <0.1 1.4 Small Grains (Prior to Creep) 47 6.6 17 19 8.4 <0.1 1 Small Grains Type I (Post Creep) 47 6.5 18 17 11 <0.1 1 Small Grains Type II (Post Creep) 45 5.7 23 15 9.8 <0.1 1.2 From Figure 56 it can be seen that the matrix in post-creep 11.6 wt% La2O3-YSZ composite hadextensive fractures along the grain boundaries. The cracks had a crenellated shape, as in the previous composites. As it was described earlier, the 11.6 wt% La2O3-YSZ composite had a bimodal microstructure composed of small and large grains. The chemical composition of the large grains did not change significantly during creep, but their size after 94 compressive creep was 2.31 μm ± 0.17 μm ( 147 % increase). In the case of the small grains, their size after creep was 0.37 μm ± 0.03 μm, which represents a 23 % decrease in size. The small grains decreased in size and also evolved two distinct chemistries as shown in Table 25. SEM observation (Figure 57) revealed that some of these small grains contained new boundaries. As a result, the small grains (after creep) have been classified into Type I and Type II. Figure 57 SEM micrographs of small grains in post-creep 11.6 wt% La2O3-YSZ: a) type I (without substructure), b) type II (with substructure). Type I grains appeared to be larger without internal boundaries, while grains with definite internal boundaries were called type II. Type I small grains (Figure 57a) had a chemistry almost identical to the small grains in the material prior to creep. However, type II small grains (Figure 57bFigure 57 ) had a slightly lower Zr, Y and O concentrations and a higher La concentration than the type I grains. The Y and La ions have comparable size and hence the difference in composition could be representative of La 3+ replacing the Y 3+ in the lattice, while Zr 4+ could be diffusing out into the rest of the matrix. With the main difference between the two types of small grains being their chemistry, it is necessary to perform XRD 95 analysis to verify possible phase changes and their stoichiometry (possibly distinct La2O3 phases or secondary phases created from the combination of ZrO2 or Y2O3). The microstructure of the composite after creep exhibited high levels of grain boundary cracking as seen in Figure 58a. In addition to the cracked boundaries, the high magnification micrographs in Figure 58 also show voids at the grain boundaries and triple points of the large grains. This suggests that the grain boundaries without La-rich grains (type I or II) often separated. Since the stress exponents was close to a value of 2, then grain boundary sliding is suspected. The La-rich grains along the boundary could have behaved as obstacles for mobility of cracks. 96 Figure 58 SEM micrographs of post-creep 11.6 wt% La2O3-YSZ showing: a) grain boundary voids, b) voids (shown by arrows) located at boundaries and triple points, c) grain boundary interface with La- rich grains separated by voids. 97 The stress exponent of ~2 suggests that grain boundary sliding was the active mechanism in compressive creep of 11.6 wt% La2O3-YSZ deformed between 1300 and 1330 °C and 45-78 MPa. Unfortunately there is no published data on compressive creep of La2O3-YSZ and thus the identification of the active creep mechanism remains inconclusive. At a temperature of 1300°C, the YSZ composite had an activation energy of 757 kJ/mol while the La2O3-YSZ composite had an activation energy of 546 kJ/mol. At a temperature of 1330 °C the YSZ composite had an activation energy of 711 kJ/mol, while the La2O3-YSZ composite had an activation energy of 586 kJ/mol. These values suggest that addition of Lanthana to YSZ reduced the activation energy at both temperatures, which, in theory, could indicate that the creep resistance of the material is hindered by La2O3 addition. Although it may take less energy to initiate deformation in YSZ when Lanthanum is present in the composite, the values for stress exponent indicate that the active mechanisms of creep (grain boundary sliding and possibly lattice diffusion) are likely the same with or without Lanthanum in the material. There is also a lack of data on the activation energy values for different creep mechanisms in La2O3-YSZ, thus the measured Q in this research cannot be compared with other studies. 98 Chapter 5: Conclusions The chapter presents the conclusions obtained from the results of the performed experiments. In this thesis, the development, characterization and compressive creep testing of rare earth oxide-doped YSZ composites processed through Spark Plasma Sintering was carried out. The findings of this research are categorized into processing, microstructure and compressive creep results. 5.1 SPS Processing  Black coloration was observed in all of the as-sintered composites and was attributed to carbon pickup from the graphite sheets used in SPS  Orange coloration was observed in the CeO2-YSZ composites and was attributed to reduction of the Cerium ion  Nearly all of the composites presented some degree of fracture; these fractures could be result from internal stresses in the material due to phase transformations occurring in the materials  Precursors powders tended to form agglomerates, which made homogeneous blending of the powders difficult. Consequently, agglomerates were observed in the microstructure after sintering  Relative densities of >95 % were obtained in 8 out of 12 of the prepared composites 5.2 Microstructure Characterization The findings related to material characterization and microstructure of YSZ, 6.4, 12.7 and 18.8 wt% CeO2-YSZ, 11.6, 21.6 and 30.5 wt% La2O3-YSZ, and 12 and 22.2 wt% Nd2O3-YSZ are:  The cubic phase of Zirconia was stabilized in all of the composites  Impurity Hf and Cl were found in the precursor powders as well as as-sintered composites. These likely originated from the fabrication of the raw powders  The increase of rare earth oxide concentration did not have a clear causal effect on the grain size or relative densities of the materials 99  Increasing the processing temperature from 1400-1500 °C resulted in grain size growth of the matrix by ~289 % in YSZ and ~425 % in La2O3-YSZ, but a reduction of ~8 % in the CeO2-YSZ composites  The average measured grain size in the general matrix of the YSZ composites was 12 µm. In the case of CeO2-YSZ composites it was 1.4 µm. For La2O3-YSZ composites it was 7.7 µm. Finally for Nd2O3-YSZ composites it was 2.6 µm  Large particles with enriched concentration of CeO2 and La2O3 were found in the microstructure of CeO2-YSZ and La2O3-YSZ composites, respectively. These particles were composed of numerous powder agglomerates that did not blend properly with the rest of the matrix powders  Addition of La2O3 and Nd2O3 to YSZ resulted in a bimodal microstructure at all doping concentrations  A secondary matrix was observed in the microstructure of La2O3-YSZ composite when the La2O3 content reached 30.5 wt% 5.3 Creep Performance The findings related to the compressive creep testing of YSZ, 6.4 wt% CeO2-YSZ and 11.6 wt% La2O3-YSZ composites are:  All of the post-creep composites showed extensive cracking along the grain boundaries as well as cavitation at grain triple points  The calculated values for n and Q in YSZ ranged from 1.7 to 2.7 and 711 to 757 kJ/mol, respectively  Unidentified round-like particles were observed throughout the YSZ microstructure, possibly consisting of tetragonal or monoclinic phase of ZrO2  The calculated values for n and Q in 6.4 wt% CeO2-YSZ composite ranged from 1.2 to 2.7 and 411 to 572 kJ/mol, respectively  The post-creep grain size in 6.4 wt% CeO2-YSZ was measured to be 3.71 μm, which represents a 147 % increase compared to the grain size before creep testing  The microstructure in post-creep 6.4 wt% CeO2-YSZ had scattered Cerium-rich particles; in cases where Zirconium diffused into those particles, the particles generally showed cracks and detachment from the matrix, while particles with much 100 lower Zirconium content did not crack and the cracks propagated around such particles  The calculated values for n and Q in the 11.6 wt% La2O3-YSZ composite ranged from 1.7 to 2.1 and 546 to 586 kJ/mol, respectively  The post-creep grain size in the 11.6 wt% La2O3-YSZ composite was 2.31 μm, which represents a 23 % increase compared to the grain size before deformation  Two distinct types of small grains having two types of chemical compositions were observed in the post-creep 11.6 wt% La2O3-YSZ composite, one type having a chemistry almost identical to the grains before creep and the other type having slightly higher La and lower Zr and Y contents  From the calculated n and Q values and the microstructural observations, the suggested creep-controlling mechanisms in all of the composites were grain boundary sliding possibly aided by simultaneous lattice diffusion 101 Chapter 6: Suggested Future Work This chapter provides suggestions for future work. This section summarizes some key areas for future research that could enable more optimal processing conditions and a better understanding of the microstructure of rare earth-doped cubic Zirconia. 6.1 Optimization of SPS-processed Zirconia Composites The recommendations for future research to achieve optimal conditions of SPS- processed rare earth-doped cubic Zirconia are as follows:  Study blending of precursor powders to disperse agglomerations; potential changes could include crushing of powders or use of milling balls  Enhance purity of precursor powders, which could be achieved by fast firing of the powders before blending  Study the effect of sintering die dimensions on densification  Use deflocculants or sintering aids to reduce aggregation  Explore alternative techniques for ejection of as-sintered materials from die to prevent fractures 6.2 Characterization of Rare Earth-doped Zirconia Composites Suggested future research areas aimed at improving the understanding of microstructure of SPS-processed rare earth-doped cubic Zirconia before and after creep:  Development of tertiary phase diagrams for the investigated materials: ZrO2-Y2O3- CeO2, ZrO2-Y2O3-La2O3, and ZrO2-Y2O3-Nd2O3, in order to identify the solubility limits of each composite and the present phases  Perform repeated and systematic sintering of the studied composites as well as additional doping concentrations and processing conditions to find any trends regarding grain size, relative density and porosity of the materials as a function of dopant concentration  Carry out extensive XRD analysis of post-creep microstructures to identify any stress or temperature induced phase transformations or precipitates 102  Conduct diffusion studies in order to confirm the suggested trends regarding solute drag and to verify the possibility of lattice diffusion as a creep-controlling mechanism  Perform TEM analysis to observe dislocation motion and attain a better understanding on the potential strain hardening, pinning or dislocation interactions happening, which would also facilitate the interpretation of the calculated n and Q values  Perform additional compressive creep testing to generate enough data to estimate the n, Q, and p (grain size exponent) 103 References [1] S. 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Verrall, “Diffusion-accomodated flow and superplasticity,” Acta Metallurgica, vol. 21, no. 1973, pp. 149–163, 1973. [76] S. Nazarpour, C. López-Gándara, F. M. Ramos, and A. Cirera, “Phase transformation studies on YSZ doped with Alumina part 1: metastable phases,” Journal of Alloys and Compounds, vol. 505, no. 2010, pp. 527–533, Sep. 2010. [77] J. A. Allemann, B. Michel, H.-B. Märki, L. J. Gauckler, and E. M. Moser, “Grain growth of differently doped Zirconia,” Journal of the European Ceramic Society, vol. 15, no. 10, pp. 951–958, Jan. 1995. 110 Appendices This section provides additional data pertaining specific chapters of this thesis. Appendix A: Sample Calculations for n and Q This Appendix shows an example of the calculations followed in order to obtain values for the stress exponent and activation energy from the varying load and temperature compressive creep tests of YSZ. The data presented here is only the information necessary to calculate one value of n and Q. A.1 Setup The calculations performed are be described in the following manner: firstly, the known constant are listed; secondly, the variables, which were measured directly from both the paper register as well as the LVDT are explained; thirdly, the necessary calculations for those variables changing with time are described; finally, the evaluation of the stress exponent and activation energy from the presented data are shown. The relevant equations used are Equations 4, 8-12, which were introduced in Chapter 3. A.2 Variables Obtained from Testing Parameters and Constants The testing parameters as well as the initial dimensions of the tested sample were known. Thus the known constants were:  testing temperature T,  load F,  initial length Lo  initial cross-section So.  gas constant R The compressive test was carried out between 1300 and 1330 °C and loads between 78 and 193 MPa, therefore variables T and F were known. The sample tested had an initial length of 4.93 mm and initial cross-section of 4.2785 mm 2 . The other constant known is the gas constant R, which has a value of 8.32 J/mol -1 K -1 . 111 A.3 Variables Obtained From Direct Measurements Directly measured variables are those that the user obtains from either the register data or the LVDT. An example of a plot obtained from the register was shown in Figure 59. This particular segment corresponds to a temperature of 1300 °C and a load of 64 kg. The slope of this line is necessary to calculate the strain rate (see Equation 9). The scale in Figure 59 is used to estimate the slope of this line in terms of micrometers per minute. Therefore the slope of the line in Figure 59 was calculated as follows: = 0.01857 Equation 13 These values of Y and t for this particular line can be seen under the columns of each variable (bolded) in the excel sheet shown in Figure 60. The table contains part of the data generated in the compressive creep test for YSZ, which corresponds to the segment where the stress was ~133 MPa in Figure 45. This column contains multiple values for Y and t because the generated line for this entire segment was divided into multiple lines (due to limitations of the register’s paper), so for each line produced for that one segment the corresponding slope was measured. For the change in length of the sample ∆L were obtained from the LVDT and were also shown in Figure 60. A.4 Variables Obtained From Calculation The variables that changes with time were:  the instantaneous length Li,  cross-section Si,  true stress i,  true strain i and  true strain rate ̇. The instantaneous length of the sample was calculated from the initial length Lo and the change in length ∆L as follows:  –   Equation 14 112 Note that the units used were micrometers as measured using the LVDT. The next calculation performed was for the instantaneous stress εi through Equation 12: ( ) (   ) Equation 15 Similarly, the instantaneous cross-section Si was calculated from Equation 11: Equation 16 The true stress i was obtained from Equation 10: ⁄ Equation 17 Finally the strain rate ̇ evaluated from Equation 9: ̇   Equation 18 The calculated values were used to evaluate the stress exponent n and activation energy Q. A.5 Calculating n and Q As it was explained in Chapter 3, the stress exponent was obtained from the data generated from one load jump and the activation energy was obtained from one temperature jump. As seen from Equation 4, the strain rates and instantaneous stress between two load jumps were needed to calculate a value of n. This means that in the columns for instantaneous strain rate and stress in Figure 60, the numbers used in the calculations are the last set of values right before a load (if calculating n) or temperature jump (if calculating Q); such values for instantaneous strain rate and stress are bolded in Figure 60under the “Stress” and “Strain Rate” columns. The calculation of n is done using Equation 4: 113 ( ̇ ) ̇⁄ ( )⁄ ( ) ( ) Equation 19 The calculated value for n was 1.877 and it is shown bolded in Figure 61 under the “Stress Exponent” column. The final calculation necessary was for Q, which was done using Equation 8: ( ( ̇) ( ̇) ) [ ( ) ( ) – ( ) ] Equation 20 The calculated value for Q is 757 kJ/mol and this value is bolded in Figure 61 under “Activation Energy”. The entire data sheet is included in Figure 62 to Figure 64. 114 Figure 59 Displacement of pistons vs. time plot obtained from register for compressive creep of YSZ at 1300 °C and 627 N. 115 Figure 60 MS Excel spreadsheet with measured and calculated variables obtained during compressive creep test of YSZ at 1300 °C and 627 N (Part 1 of 4). 116 Figure 61 MS Excel sheet with calculated values for n and Q from a load and temperature jump. 117 Figure 62 MS Excel spreadsheet with measured and calculated variables obtained during compressive creep test of YSZ at 1300 °C and 627 N (part 2 of 4). 118 Figure 63 MS Excel spreadsheet with the measured and calculated variables obtained during compressive creep test of YSZ at 1300 °C and 627 N (part 3 of 4). 119 Figure 64 MS Excel spreadsheet with the measured and calculated variables obtained during compressive creep test of YSZ at 1300 °C and 627 N (part 4 of 4). 120 Appendix B: Micrographs Figure 65 SEM-XEDS line-scan analysis of Ce-rich particle in 12.7 wt% CeO2-YSZ composite. 121 Figure 66 SEM-XEDS line-scan analysis of Ce-rich particle in 18.8 wt% CeO2-YSZ composite. 122 Figure 67 SEM micrograph of small grains in 12 wt% Nd2O3 – YSZ at 100,000x magnification.
Thesis/Dissertation
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Mechanical Engineering
Vancouver : University of British Columbia Library
Attribution-NonCommercial-NoDerivatives 4.0 International
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Development, characterization and creep resistance of cubic zirconia doped with rare earth oxides processed through spark plasma sintering
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