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Low-Voltage, Large-Range MEMS Optical Scanners and Their Applications

Permanent Link: http://ufdc.ufl.edu/UFE0024705/00001

Material Information

Title: Low-Voltage, Large-Range MEMS Optical Scanners and Their Applications
Physical Description: 1 online resource (191 p.)
Language: english
Creator: Wu, Lei
Publisher: University of Florida
Place of Publication: Gainesville, Fla.
Publication Date: 2009

Subjects

Subjects / Keywords: aligned, aperture, array, bimorph, biomedical, circumferential, coherence, communications, confocal, deep, displacement, dual, electrothermal, endoscope, folded, fourier, free, full, imaging, insensitive, large, lateral, light, low, microelectromechanical, microlens, micromirror, microscopy, miniature, mirror, nonlinear, optical, phased, piston, range, reflective, scanner, self, shift, space, spectrometer, system, tilt, tip, tomography, transform, trench, vertical, voltage, weight
Electrical and Computer Engineering -- Dissertations, Academic -- UF
Genre: Electrical and Computer Engineering thesis, Ph.D.
bibliography   ( marcgt )
theses   ( marcgt )
government publication (state, provincial, terriorial, dependent)   ( marcgt )
born-digital   ( sobekcm )
Electronic Thesis or Dissertation

Notes

Abstract: Advanced imaging techniques including optical coherence tomography/microscopy (OCT/OCM), nonlinear optical (NLO) imaging and confocal microscopy demonstrate powerful resolution and optical-sectioning capabilities; therefore can potentially replace conventional biopsy diagnosis procedures for early cancer detections. To realize in vivo, noninvasive clinical endoscopic imaging, miniature endoscopes integrated with small, versatile and large range optical scanners including 1-D, 2-D transverse and full-circumferential scan micromirrors, as well as large-axial-scan microlens scanners must be developed. The objective of this research is to develop miniature optical scanners by Microelectromechacnial Systems (MEMS) technology and the MEMS-based in vivo biomedical imaging endoscopes. Several novel actuators based on electrothermal bimorph actuation are developed in this work that solve problems in previous generations including large mirror center shift, large initial tilting and elevation, complicated mirror control, and low fill factor. The lateral-shift-free (LSF) large-vertical-displacement (LVD) actuator realizes versatile optical scanners including tip-tilt-piston (TTP) mirrors, lens scanners and large-aperture mirrors with large axial scan. The TTP mirror demonstrates 2-D tip-tilt scan > 60masculine ordinal and piston scan > 0.6 mm at < 5 Vdc with an improved fill-factor of 25 %. Over 0.9 mm large axial scan mirrors and lens scanners with small tilting below 0.4masculine ordinal are also presented. The dual-folded-bimorph (DFB) actuator realizes over 90masculine ordinal mechanical rotation up to 60 Hz with a stationary center rotation axis and a flat, un-elevated initial mirror position, full-circumferential scan at real time imaging speed is achieved by the DFB-based dual-reflective micromirror. A novel self-aligned deep-trench process is also developed to fabricate the dual-reflective miromirror and light-weight large-aperture mirrors. MEMS imaging endoscopes for both OCT and NLO are developed; 3-D in vivo imaging results are successfully demonstrated. Other potential applications are also investigated. A 4times4 TTP mirror array with sub-aperture size of 0.9 mm and a fill-factor of 65% is presented for optical phased array application. MEMS mirror with large aperture up to 10 mm, tip-tilt scan of ~10masculine ordinal and resonance in the order of 100 Hz are demonstrated for free-space optical communications. The prototypes of miniature Fourier transform spectrometer (FTS) are demonstrated, the large-axial-scan MEMS mirror combined with a novel mirror-tilt-insensitive FTS system has achieved a high spectral resolution of 8.1 cm-1.
General Note: In the series University of Florida Digital Collections.
General Note: Includes vita.
Bibliography: Includes bibliographical references.
Source of Description: Description based on online resource; title from PDF title page.
Source of Description: This bibliographic record is available under the Creative Commons CC0 public domain dedication. The University of Florida Libraries, as creator of this bibliographic record, has waived all rights to it worldwide under copyright law, including all related and neighboring rights, to the extent allowed by law.
Statement of Responsibility: by Lei Wu.
Thesis: Thesis (Ph.D.)--University of Florida, 2009.
Local: Adviser: Xie, Huikai.
Electronic Access: RESTRICTED TO UF STUDENTS, STAFF, FACULTY, AND ON-CAMPUS USE UNTIL 2010-02-28

Record Information

Source Institution: UFRGP
Rights Management: Applicable rights reserved.
Classification: lcc - LD1780 2009
System ID: UFE0024705:00001

Permanent Link: http://ufdc.ufl.edu/UFE0024705/00001

Material Information

Title: Low-Voltage, Large-Range MEMS Optical Scanners and Their Applications
Physical Description: 1 online resource (191 p.)
Language: english
Creator: Wu, Lei
Publisher: University of Florida
Place of Publication: Gainesville, Fla.
Publication Date: 2009

Subjects

Subjects / Keywords: aligned, aperture, array, bimorph, biomedical, circumferential, coherence, communications, confocal, deep, displacement, dual, electrothermal, endoscope, folded, fourier, free, full, imaging, insensitive, large, lateral, light, low, microelectromechanical, microlens, micromirror, microscopy, miniature, mirror, nonlinear, optical, phased, piston, range, reflective, scanner, self, shift, space, spectrometer, system, tilt, tip, tomography, transform, trench, vertical, voltage, weight
Electrical and Computer Engineering -- Dissertations, Academic -- UF
Genre: Electrical and Computer Engineering thesis, Ph.D.
bibliography   ( marcgt )
theses   ( marcgt )
government publication (state, provincial, terriorial, dependent)   ( marcgt )
born-digital   ( sobekcm )
Electronic Thesis or Dissertation

Notes

Abstract: Advanced imaging techniques including optical coherence tomography/microscopy (OCT/OCM), nonlinear optical (NLO) imaging and confocal microscopy demonstrate powerful resolution and optical-sectioning capabilities; therefore can potentially replace conventional biopsy diagnosis procedures for early cancer detections. To realize in vivo, noninvasive clinical endoscopic imaging, miniature endoscopes integrated with small, versatile and large range optical scanners including 1-D, 2-D transverse and full-circumferential scan micromirrors, as well as large-axial-scan microlens scanners must be developed. The objective of this research is to develop miniature optical scanners by Microelectromechacnial Systems (MEMS) technology and the MEMS-based in vivo biomedical imaging endoscopes. Several novel actuators based on electrothermal bimorph actuation are developed in this work that solve problems in previous generations including large mirror center shift, large initial tilting and elevation, complicated mirror control, and low fill factor. The lateral-shift-free (LSF) large-vertical-displacement (LVD) actuator realizes versatile optical scanners including tip-tilt-piston (TTP) mirrors, lens scanners and large-aperture mirrors with large axial scan. The TTP mirror demonstrates 2-D tip-tilt scan > 60masculine ordinal and piston scan > 0.6 mm at < 5 Vdc with an improved fill-factor of 25 %. Over 0.9 mm large axial scan mirrors and lens scanners with small tilting below 0.4masculine ordinal are also presented. The dual-folded-bimorph (DFB) actuator realizes over 90masculine ordinal mechanical rotation up to 60 Hz with a stationary center rotation axis and a flat, un-elevated initial mirror position, full-circumferential scan at real time imaging speed is achieved by the DFB-based dual-reflective micromirror. A novel self-aligned deep-trench process is also developed to fabricate the dual-reflective miromirror and light-weight large-aperture mirrors. MEMS imaging endoscopes for both OCT and NLO are developed; 3-D in vivo imaging results are successfully demonstrated. Other potential applications are also investigated. A 4times4 TTP mirror array with sub-aperture size of 0.9 mm and a fill-factor of 65% is presented for optical phased array application. MEMS mirror with large aperture up to 10 mm, tip-tilt scan of ~10masculine ordinal and resonance in the order of 100 Hz are demonstrated for free-space optical communications. The prototypes of miniature Fourier transform spectrometer (FTS) are demonstrated, the large-axial-scan MEMS mirror combined with a novel mirror-tilt-insensitive FTS system has achieved a high spectral resolution of 8.1 cm-1.
General Note: In the series University of Florida Digital Collections.
General Note: Includes vita.
Bibliography: Includes bibliographical references.
Source of Description: Description based on online resource; title from PDF title page.
Source of Description: This bibliographic record is available under the Creative Commons CC0 public domain dedication. The University of Florida Libraries, as creator of this bibliographic record, has waived all rights to it worldwide under copyright law, including all related and neighboring rights, to the extent allowed by law.
Statement of Responsibility: by Lei Wu.
Thesis: Thesis (Ph.D.)--University of Florida, 2009.
Local: Adviser: Xie, Huikai.
Electronic Access: RESTRICTED TO UF STUDENTS, STAFF, FACULTY, AND ON-CAMPUS USE UNTIL 2010-02-28

Record Information

Source Institution: UFRGP
Rights Management: Applicable rights reserved.
Classification: lcc - LD1780 2009
System ID: UFE0024705:00001


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1 LOW -VOLTAGE, LARGE RANGE MEMS OPTICAL SCANNERS AND THEIR APPLICATIONS By LEI WU A DISSERTATION PRESENTED TO THE GRADUATE SCHOOL OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF DOCT OR OF PHILOSOPHY UNIVERSITY OF FLORIDA 2009

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2 2009 Lei Wu

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3 To my parents, Jinlong Wu and Xinglan Zhou to m y wife Qi Zhou and t o my daughter Meining

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4 ACKNOWLEDGMENTS I would like to gratefully and sincerely thank my advisor, Professor Huikai Xie for his guidance, encouragement, friendship, and all the opportunities and support that he offered me during the past four years of my Ph.D. study and research Dr. Xies mentor ship wa s well rounded with his profound knowledge and extensive experience in M icroelectromechanical Systems (MEMS) his enlightenment and encouragement on my creativity, and with his training that cultivates me towards an all rounded MEMS engineer. I was truly given an open, free and well -supported environment to conceive new ideas and conduct my research. And, most excitingly, upon the completion of this dissertation, I was offered the great opportunity to join his start up company and continue to work as his colleague to promote our research accomplishments in to the real world. T hanks to Professor s Peter Zory, Toshikazu Nishida and Huabei Jiang for serving on my dissertation committee Dr. Zory is one of the greatest scientist and experimentalist that I ve ever met. I had a great opportunity assisting Dr. Zorys semiconductor laser research for one semester, and was greatly impresse d by his rigorous and meticulous attitude towards the scientific experiments, which I believe will continuously inspire me in my future career. I am also grateful for my personal friendship with Dr. Zory and his generosity for constantly supporting our research with his lab facilities. I thank Dr. Nishida for his insightful suggestions from the proposal of this research. As an M EMS specialist, Dr. Nishidas questions are always challenging, and I enjoyed many of the enlightening discussions with him. It was also my pleasure to serve in the first Interdisciplinary Microsystems Group (IMG) safety committee that is led by Dr. Nishid a and I saw a group -wide mentor who is caring, resp onsible and dedicated I am grateful to Dr. Jiangs valuable suggestions as a biomedical imaging specialist, and the newly

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5 collaborated project that will use some of my devices. I believe that more resear ch accomplishments will be achieved very soon. My acknowledgement goes to Dr. Shuguang Guo, a great o ptical c oherence t omography (OCT ) specialist. He joined the lab at the right time when my devices are ready for OCT imaging experiments. He spent just one month to build the current OCT system at B iophotonics and Microsystems Lab (BML), and continuously volunteered himself as the sample for many of our in vivo OCT imaging experiments using my MEMS endoscopes. As an opticist, he also contributed his knowl edge and creativity in our MEMS based Fourier t ransform s pectrometer (FTS) project and helped in many of my optical experiments. Thanks to the other OCT team members at BML, Jingjing Sun and Lin Liu, who helped on the OCT endoscope design, assembly and imagin g experiments. B oth of them proved to be qualified successors of Dr. Guo, Jingjing is more focusing on the endoscopic OCT and Lin is taking over some of my devices and exploring the endoscopic confocal imaging and full -circumferential scanning OCT. I wish great success for both of them. Special thanks to the other members of FTS team Andrea Pais and Sean R. Samuelson for their contributions in conceiving our novel system and efforts in data collection and processing, and for our unforgettable friendships We spent many nights working hard to catch up the conference deadline and I still remember all the jokes we had that helped us to survive the exhausting experiments. I thank our external colla borator the nonlinear optical imaging specialists at Swinburne University of Technology Australia, Professor Min Gu, Dr. Lin Fu, and especially Dr. Dru Morrish, for his time and efforts dedicated to this project and our personal friendship. As one of the second batch Ph.D s in BML, I had an overlap of one year with the first batch Ph.D.s in the lab, Dr. Hongwei Qu, Dr. Ankur Jain and Dr. Deyou Fang who, as helpful

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6 and caring senior students, gave me the technical trainings for lab safety and cleanroom facilities and helped me to survive the first year in the lab. I also benefited a lot from the professional interactions with them and am grateful to their friendships. I value the camaraderie and thank the help from other BML members including Mingliang Wang, Kemiao Jia, Sagnik Pal Hongzhi Sun and Dr. Yiping Zhu Kem iao was also a lab mate in my former institute; we joined the lab at the same time and shared the stresses by helping each other when we first started our research I enjoyed working with Dr. Zhu to expand my electrothermal actuator design to piezoelectric actuation, discussing with Sagnik about modeling issues sharing the fabrication recipes with Mingliang and seeking advice from Hongzhi about circuit issues. I will treasure all my memories with the BML family, in the Larsen 136 office, in the cleanroom, in the group meetings, as well as the sports and all the group parties that were generously sponsored by Dr. Xie. It is really exciting to see that BML has grown into a big research lab with nine Ph.D. and five master students, as well as postdocs and visi ting scholars. I believe that greater success of BML will be continuously achieved. IM G is a bigger family of MEMS researchers at University of Florida. I have been taking the advantages of this joint group including the easy access to all the shared faci lities and equipments, the resource of technical expertise, and most importantly, the active research and experienced researchers in diverse MEMS area Thanks to all IMG faculties and other members incl u d ing Dr. Jian Liu, Dr. Benjamin Griffin, Alex Phipps, Tai -An Chen, Dylan Alexander Brandon Bertolucci Jeremy Sells and whoever aided in my experiments and provided technical support All the MEMS devices in this dissertation were fabricated in the former University of Florida Nanofabrication Facilities (U FNF), which has turned to the current Nanoscale Research

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7 Facility (NRF). I would like to acknowledge the facilit y maintenance and technical support provided by the NRF staffs, Bill Lewis, Ivan Kravchenko, and especially Al Ogden, who never refused to come and help when I met problems in the Benton cleanroom during after hours I also thank the Major Analytical Instrumentation Center (MAIC) where the scanning electron microscopy (SEM) and white light profilometry of the MEMS devices have been performed. La st but never the least, my deepest gratitude goes to my fami ly. I thank my parents, Jinlong Wu and Xinglan Zhou who made all of this possible, for their endless love, support, encouragement, patience, and for their instilling of diligence, perseverance an d confidence. I thank my wife Qi Zhou, who has been proud and supportive of my work, for her constant love, care, understanding and sharing of all the uncertainties, challenges and sacrifices in the past four years of our life She has been alone for count less nights and weeken ds while I was working in the lab, I promise that I ll spend much more time with her from now on. And an additional huge thank you to her for the most fantastic gift our daughter Meining, who was born just two weeks before the comple tion of this dissertation. I am also grateful to my parents in law, who came all the way from Shanghai to help with the baby -sitting so that I can focus on this dissertation during the last few weeks before my defense. This research has been supported by National Science Foundation under Award # 0423557, 0725598, 0818473, Kodak, and Florida Photonics Center of Excellence.

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8 TABLE OF CONTENTS page ACKNOWLEDGMENTS .................................................................................................................... 4 LIST OF TABLES .............................................................................................................................. 11 LIST OF FIGURES ............................................................................................................................ 12 ABSTRACT ........................................................................................................................................ 17 CHAPTER 1 INTRODUCTION ....................................................................................................................... 19 1.1 Conventional Cancer Detection Methods and Their Limitations ...................................... 20 1.2 New Biomedical Imaging Techniques ................................................................................. 21 1.2.1 Optical Coherence Tomography ............................................................................ 21 1.2.2 Nonlinear Optical Microscopy............................................................................... 23 1.2.3 Confocal Laser Scanning Microscopy .................................................................. 24 1.3 MEMS -based Endoscope ..................................................................................................... 26 1.4 Research Goal and Tasks ...................................................................................................... 28 1.5 Dissertation Overview .......................................................................................................... 29 2 MEMS ELECTROTHERMAL OPTICAL SCANNERS ......................................................... 32 2.1 MEMS Optical Scanners ...................................................................................................... 32 2.2 Principle of Electrothermal Bimorph Actuation ................................................................. 37 2.2.1 Stress and Curvature in Bimorph Cantilever ........................................................ 38 2.2.2 Bimorph Sensitivity and Material Selection ......................................................... 41 2.2.3 Electrothermal Analysis ......................................................................................... 43 2.3 Prior Electrothermal Actuator Designs ................................................................................ 45 2.4 Summary of Remaining Problems ....................................................................................... 48 3 LATERAL SHIFT FREE LARE -VE RTICAL DISPLACEMENT ELECTRO THERMAL ACTUATOR AND MICROSC ANNERS ................................................................................. 51 3.1 LSF -LVD Electrothermal Actuator Design ........................................................................ 52 3.2 LSF -LVD based Tip Tilt -Piston Micromirror .................................................................... 56 3.2.1 Device Design ......................................................................................................... 56 3.2.2 Device Fabrication .................................................................................................. 59 3.2.3 Experimental Results .............................................................................................. 61 3.2.3.1 Piston motion................................................................................................ 6 1 3.2.3.2 Tip -tilt motion .............................................................................................. 65 3.2.3.3 Dynamic response ........................................................................................ 67 3.3 LSF -LVD Microlens Scanner .............................................................................................. 69

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9 3.3.1 Device Design and Fabri cation .............................................................................. 69 3.3.2 Experimental Results .............................................................................................. 71 3.4 Lumped Element Models and Improved LSF LVD Actuator Design .............................. 72 3.4.1 SteadyState Analysis ............................................................................................. 73 3.4.2 Transient Thermal Analysis ................................................................................... 75 3.4.3 Improved LSF -LVD Actuator Design with Faster Thermal Response ............... 78 3.4.3.1 Reducing the Thermal Resistance ............................................................... 78 3.4.3.2 Reducing the Thermal Capa citance ............................................................ 80 3.5 Summary ............................................................................................................................... 88 4 FULL -CICURMFERENTIAL SCAN NING MICROMIRRORS AN D LIGHT WEIGHT MICROMIRRORS ...................................................................................................................... 89 4.1 Full Circumferential Scanning Endoscope Design ............................................................ 89 4.2 Dual Reflective Micromirror ............................................................................................... 91 4.2.1 Device Design ......................................................................................................... 91 4.2.2 Device Fabrication .................................................................................................. 93 4.2.3 SteadyState Equivalent Circuit ............................................................................. 97 4.2.4 Experimental Results ............................................................................................ 100 4.2.4.1 Static response ............................................................................................ 100 4.2.4.2 Dynamic response and circumferen tial scan ............................................ 102 4.3 Dual Folded -Bimorph Actuator and FCS Micromirrors .................................................. 104 4.3.1 Dual Folded -Bimorph Actuator Design .............................................................. 105 4.3.2 Device Fabrication ................................................................................................ 106 4.3.3 Lumped Element Model ....................................................................................... 108 4.3.4 Device C haracterization ....................................................................................... 112 4.4 Light -Weight Micromirrors ............................................................................................... 116 4.4.1 Self -Aligned Deep Trench Process ..................................................................... 117 4.4.2 Experimental Results ............................................................................................ 120 4.5 Summary ............................................................................................................................. 123 5 APPLICATIONS OF MEMS OPTICAL SCANNERS .......................................................... 124 5.1 MEMS -based Nonlinear Optical Endoscope .................................................................... 124 5.1.1 Nonlinear Optical Imaging System ..................................................................... 125 5.1.2 Nonlinear Optical Endoscope Design and Imaging Results .............................. 127 5.2 MEMS -based Optical Coherence Tomography Endoscope ............................................ 130 5.2.1 OCT System .......................................................................................................... 131 5.2.2 OCT Endoscope Design and Imaging Results .................................................... 132 5.3 MEMS Microlens Scanners ............................................................................................... 135 5.3.1 New LSF LVD Lens Scanner Design and Fabrication ...................................... 136 5.3.2 Experimental Results ............................................................................................ 138 5.4 MEMS Optical Phased Arrays ........................................................................................... 140 5.4.1 Mirror Array Design and Fabrication .................................................................. 142 5.4.2 Device Characterization ....................................................................................... 144 5.4.3 Phase Control Experiment ................................................................................... 148 5.5 Large -Aperture MEMS Mirror for Free Space Optical Communications ...................... 150

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10 5.5.1 Large -Aperture Mirror Design and Fabrication ................................................. 151 5.5.2 Device Characterization ....................................................................................... 152 5.6 MEMS -based Miniature Fourier Transform Spectrometer .............................................. 155 5.6.1 Fourier Transform Spectrometer (FTS) System ................................................. 155 5.6. 2 MEMS -based Miniature FTS ............................................................................... 157 5.6.3 LSF -LVD MEMS mirror for Miniature FTS ...................................................... 158 5.6.4 Device Characterization and FTS Experiment ................................................... 159 5.6.5 Mirror Tilt Insensitive FTS System .................................................................... 163 5.6.5.1 MTI -FTS setup ........................................................................................... 163 5.6.5.2 Experimental results .................................................................................. 165 5.7 Summary ............................................................................................................................. 167 6 CONCLUSION AND FUTUR E WORK ................................................................................ 169 6.1 Research Accomplishments ............................................................................................... 170 6.2 Future Work ........................................................................................................................ 171 APPENDIX PUBLICATIONS AND PRO VISONAL PATEN TS GENERATED BY THIS RESEARCH EFFORT .............................................................................................................. 172 LIST OF REFERENCES ................................................................................................................. 175 BIOGRAPHICAL SKETCH ........................................................................................................... 191

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11 LIST OF TABLES Table page 2 1 Thermomechanical properties (at room temperature) of materials possible for bimorph actuation ................................................................................................................... 43 3 1 FEM simulation results of TTP micromirror design by Coventorware .............................. 58 3 2 Structure parameters of fabricated TTP micromirrors ......................................................... 60 3 3 Measured piston resonance, DC vertical displacements and estimated average temperature on the bimorph at different DC bias, for a modified Type II device ............. 68 3 4 Parameters used in the lumped element model LSF -LVD actuator .................................... 76 3 5 Material and structure parameters used for the analytical calculation ................................ 84 3 6 Summar y of the transient performance of the device Type II III and IV ......................... 87 4 1 Parameters used in the steady -state lumped element model for 1 -D micromirror ............. 98 4 2 Parameters used in the lumped element models for the DFB -FCS mirror ....................... 109 4 3 Comparison of resonance frequencies of complete-Si device and light -weight devices 123 5 1 Glass lens parameters (1300nm wavelength). .................................................................... 138 5 2 Parameters of the fabricated mirror arrays ......................................................................... 144 5 3 MEMS mirror scan ranges, tilting angles and spectral resolutions at different driving voltages ................................................................................................................................. 163

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12 LIST OF FIGURES Figure page 1 1 Simplified schematic of time -domain OCT setup ................................................................ 22 1 2 Simplified schematic for TPEF and SHG imaging setup .................................................... 24 1 3 Simplified schematic of a fluorescence confocal microscopy ............................................ 26 2 1 A side view of the bimorph beams with initial upward curling .......................................... 40 2 2 Schematic of 1 D micromirror design illustrating electrothermal model ........................... 44 2 3 Electrothermal micromirror design and SEM pictures of fabricated devices .................... 46 2 4 LVD electrothermal micromirror design and SEM pictures of fabricated devices. .......... 47 2 5 Cross -sectional view of LVD actuator showing the lateral shift ........................................ 48 3 1 3 D model of LSF -LVD actuator design with nomenclatures ............................................ 52 3 2 Cross -sectional view schematics. .......................................................................................... 53 3 3 Schematic showing geometry design of the LSF LVD actuator ......................................... 54 3 4 FEM simulation results of a LSF LVD actuators vertical displacements at different temperature by Coventorware. .............................................................................................. 56 3 5 Top view layout of LSF -LVD based TTP micromirror. ...................................................... 57 3 6 FEM simulation results of LSF LV D TTP micromirror by Coventorware ........................ 58 3 7 Cross -sectional view of fabrication process flow for a TTP micromirror .......................... 60 3 8 SEM pictu res of fabricated TTP micromirrors. .................................................................... 61 3 9 Piston motion static response of TTP micromirrors ............................................................ 63 3 10 Microsopic pictures of a Type II device at different applied voltages ............................... 63 3 11 Experimental results of Type II device. ................................................................................ 64 3 12 Zygo interfero metric measurem ent ....................................................................................... 65 3 13 Static tip -tilt angles versus applied voltages on individual actuators of TTP micromirrors ........................................................................................................................... 65 3 14 2 D scanning patt erns of TTP micromirrors ......................................................................... 66

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13 3 15 Frequency response measurement of TTP micromirrors by a laser vibrometer ................ 67 3 16 Measured pisto n resonances of a modifiedType II device at different DC bias ................ 68 3 17 Schematics of LSF LVD lens scanner design ...................................................................... 69 3 18 SEMs of fabricated LSF LVD lens holder ........................................................................... 70 3 19 Lens scanner imaging experiment ......................................................................................... 71 3 20 Le ns scanner experimental results ........................................................................................ 72 3 21 Microscopic pictures of the microlens at different applied voltages .................................. 72 3 22 Schematics showing actuation principle ............................................................................... 74 3 2 3 Applied voltage versus vertical displacement for a Type II device with experimental and calculated results by different values of equivalent thermal resistance. ...................... 75 3 2 4 The equivalent circuit of a simplified transient lumped element model of LSF LVD actuator .................................................................................................................................... 75 3 2 5 Measured frequency response of a Type II device ............................................................... 77 3 2 6 SEM picture of the modified LSF LVD actuator with additional Al layer on the substrate junction. ................................................................................................................... 79 3 2 7 Measured frequency response of a Type III device ............................................................. 79 3 2 8 The side -view schematic of the multimorph structure ......................................................... 82 3 2 9 The calculated reciprocal of curvature radi us versus the top SiO2 (PECVD) thickness ... 83 3 30 FEM simulation results of the multimorph structure ........................................................... 85 3 3 1 SEM pictures of a TT P micro mirror (device Type IV ) by the Si -free LSF LVD actuator. ................................................................................................................................... 86 3 3 2 Measured frequency response of a Type IV device ............................................................. 87 4 1 T he conceptual schematic of the FCS EOCT imaging probe with a dual reflective MEMS mirror ......................................................................................................................... 90 4 2 The schematic of the dual reflective electrothermal micromirror design .......................... 91 4 3 FEM simulation results of a hinged-bimorph 1 D mirror by CoventorWare ..................... 92 4 4 Theoretical calculation and FEM simulation results of rotational angl e versus average temperature change for a 1 D hinged-bimorph micromirror design, .................... 93

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14 4 5 Cross -sectional view of the fabrication process flow for the dual reflective micromirror ............................................................................................................................. 94 4 6 SEM pictures of fabricated 1 D hinged-bimorph micromirrors .......................................... 96 4 7 Line scans of the fabricated dual reflective mirror surfaces by a whit e light optical profilometer. ........................................................................................................................... 96 4 8 Equivalent circuit model of a 1 D hinged -bimorph and heater aside electrothermal micromirror ............................................................................................................................. 97 4 9 Voltage versus rotation angle from steady-state model calculations and an actually fabricated 1 D hinged-bimorph micromirror ........................................................................ 99 4 10 SEMs of thermal isolation region ........................................................................................ 100 4 11 Measured applied voltage versus rotation angle and heater electrical resistance of 1 D hinged -bimorph micromirrors ......................................................................................... 101 4 12 Comparison o f theoretical calculation, FEM simulation and experimental results of 1 D hinged -bimorph micromirrors ...................................................................................... 101 4 13 Circumferential scanning experiment. ................................................................................ 102 4 14 Measured frequency response of a dual reflective 1 D mirror ......................................... 103 4 15 3 D models of the dual reflective micromirror based on the DFB actuator design by IntelliSuit e ............................................................................................................................. 105 4 16 Fabrication process flow for a DFB -FCS micromirror ...................................................... 107 4 17 SEMs of fabricated DFB -FCS micromirror ....................................................................... 108 4 1 8 The LEM equivalent circuits of DFB -FCS micro mirror s .................................................. 109 4 1 9 Frequency response s of the DFB actuator predicted by the LEM .................................... 112 4 20 Measured DC and AC response of rotation angle versus voltage of a Type II DFB FCS micromirror. ................................................................................................................. 113 4 2 1 Measured frequency response of a Type II DFB FCS micromirror. ................................ 114 4 2 2 Microscopic pictures of a Type II DFB -FCS micromirror ............................................... 115 4 2 3 Full -circumferential scanning patterns at 30Hz by DFB FCS micromirrors ................... 115 4 24 Self aligned deep trench process flow for fabricating light -weight micromirror ............ 118 4 25 Layout and SEMs of the light -weight micromirrors. ......................................................... 120

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15 4 26 DC response of the piston and tip-tilt scan versus the applied voltages for a Type II light -weight micromirror. .................................................................................................... 121 4 27 Microscopic images of the Type II device mirror surface with piston actuation at different voltages. ................................................................................................................. 122 5 1 Nonlinear opti cal imaging system ....................................................................................... 125 5 2 3 D model of the nonlinear optical imaging endoscope .................................................... 127 5 3 Pictures of fabricated NLO probe ....................................................................................... 128 5 4 Experimental results of the NLO probe with MEMS mirror assembled .......................... 129 5 5 Two -photon fluorescence image stacks of fluorescent be ads, scale bars 10 m. ............ 130 5 6 Schematic of the time domain OCT system ....................................................................... 132 5 7 3 D model and pictures of the OCT probe ......................................................................... 133 5 8 2 D and 3 D in vivo OCT images of mouse tongue ........................................................... 134 5 9 In vivo images of mouse ear ................................................................................................ 135 5 10 Lens holder by modified LSF -LVD actuator design ......................................................... 136 5 11 SEMs of fabricated lens scanners ........................................................................................ 137 5 12 DC response of vertical displacement versus voltage ........................................................ 138 5 13 Pictures of lens B scanner and microscopic images of transparency masks formed by the lens at different actuation voltages ................................................................................ 139 5 14 Frequency responses of the unloaded lens holder, the lens scanners assembled with lens A and lens B. ................................................................................................................. 139 5 15 Tilting angles of a lens B scanner with and without the OVR .......................................... 140 5 16 SEM pictures of TTP mirror arrays .................................................................................... 143 5 17 Mirror surface measurement of a Typ e B device by a Wyko NT9800 white light interferometer ....................................................................................................................... 144 5 18 Piston actuation DC response of the TTP mirror arrays .................................................... 145 5 19 T ip -tilt actuation DC response of the TTP mirror arrays ................................................... 145 5 20 Frequency responses of Type B TTP mirror array ............................................................. 147

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16 5 21 Transien t responses of TTP mirror arrays .......................................................................... 147 5 22 Schematics of phase control experiments ........................................................................... 148 5 23 Intensity profiles of the diffraction patterns by phase control of the TTP mirror array .. 149 5 24 SEMs of fabricated largeaperture micromirrors ............................................................... 151 5 25 Measured DC responses of large aperture micromirrors ................................................... 153 5 26 Measured frequency responses of both 6 -mm and 10 -mm large aperture mirrors .......... 153 5 27 Pictures and intensity profiles of the reflected beams (incident beam size ~5 mm) by micromirrors with different aperture sizes ......................................................................... 154 5 28 A simplified schematic of an FTIR spectrometer .............................................................. 155 5 29 SEM pictures of LVD micromirror ..................................................................................... 159 5 30 DC response of LVD micromirror ...................................................................................... 160 5 31 Tilting angles measured by a PSD with same voltage on both actuators and minimized tilting angle at optimum voltage ratio. ............................................................. 160 5 32 MEMS -based FTS set up .................................................................................................... 161 5 33 Interferogram of a He -Ne laser by actuation voltages of 0.451.25 V and 0.451.32 V on the two actuators respectively ........................................................................................ 161 5 34 Spectr a of He Ne laser obtained by MEMS based FTS with different spectral resolutions ............................................................................................................................. 162 5 35 MTI -FTS system .................................................................................................................. 164 5 36 Dual reflective LVD MEMS mirror characterization ........................................................ 165 5 37 Interferograms of the He -Ne laser ....................................................................................... 166 5 38 Spectra of He Ne laser obtained by the MTI FTS system with MEMS mirror being actuated at different ramp waveforms ................................................................................. 167

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17 Abstract of Dissertation Presented to the Graduate School of the University of Florida in Partial Fulfillment of the Requirements for the Degree of Doctor of Philosophy LOW -VOLTAGE LARGE RANGE MEMS OPTICAL SCANNERS AND THEIR APPLICATIONS By Lei Wu August 2009 Chair: Huikai Xie Major: E lectrical and Computer Engineering A dvanced imaging techniques including optical coherence tomo graphy/microscopy (OCT/OCM), nonlinear optical (N LO ) imaging and confocal microscopy demonstrate powerful resolution and optical -sectioning capabilities; therefore can potentially replace conventional biopsy diagnosis procedures for early cancer detections T o realize in vivo noninvasive clinical endoscopic imaging, miniature endoscopes integrated with small, versatile and large range optical scanners including 1 D, 2 D transverse and full -circumferential scan micromirrors, as well as large axial -scan micr olens scanners must be developed. The objective of this research is to develop miniature optical scanners by Microelectromechacnial S ystems (MEMS) technology and the MEMS based in vivo biomedical imaging endoscopes. Several novel actuators based on electr othermal bimorph actuation are developed in this work that solve problems in previous generations including large mirror center shift, large initial tilting and elevation, complicated mirror control, and low fill factor. The lateral -shift -free (LSF) large vertical displacement (LVD) actuator realizes versatile optical scanners including tip tilt piston (TTP) mirror s lens scanners and large aperture mirrors with large axial scan. The TTP mirror demonstrate s 2 D tip tilt scan > 60 and piston scan > 0.6 mm at <5 Vdc with an improved fill -factor of 25 %. O ver 0.9 mm l a rge axial scan mirrors and lens scanners wit h small tilting

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18 below 0.4 are also presented. The dual -folded -bimorph (DFB) actuator realizes over 90 mechanical rotation up to 60 Hz with a stationary center rotation axis and a flat, un -elevated initial mirror position full -circumferential scan at real time imaging speed is achieved by the DFB -based dual reflective micro mirror. A novel self aligned deeptrench process is also developed to fabricate th e dual reflective miromirror and light -weight large -aperture mirrors. MEMS imaging endoscopes for both OCT and N LO are developed; 3 D in vivo imaging results are successfully demonstrated. Other potential applications are also investigated A 44 TTP mirro r array with sub aperture size of 0.9 mm and a fill -factor of 65% is presented for optical phased array application. MEMS mirror with large aperture up to 10 mm, tip tilt scan of ~10 and resonance in the order of 100 Hz are demonstrated for free -space opt ical communications. The prototypes of miniature Fourier transform spectrometer (FTS) are demonstrated, the large axial -scan MEMS mirror combined with a novel mirror tilt insensitive FTS system has achieved a high spectral resolution of 8.1 cm1.

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19 CHAPTER 1 INTRODUCTION Cancer continues to be one of the top killer diseases worldwide and accounts for one of every four deaths in U.S The estimated new cancer cases and deaths in US alone for the year 2008 are about 1,437,180 and 565, 650 respectively [1]. The death rate of cancer has declined averagely 1 .1 percent per year from 1993 to 2002, and 2.1 percent per year from 2002 to 2004 thanks to the efforts on prevention, treatment, and especially, early detections [2]. The American Cancer Society estimates that relative survival for most cancers such as breast, tongue, mouth, colon, rectum, cervix, prostate, testis and melanoma cancers would drastically incre ase if they were detected at their early stage s [1]. Recently developed imag ing techniques such as optical coherence t omogr aphy (OCT) [3 5] n on linear optical imaging (such as two -photon excitation fluorescence and second harmaonic generation) [6 9] and confocal m icroscopy [10 12] demonstrated superior performances including ultra -high resolution s up to micrometers or even sub-micrometer an d millimeters range of penetration depth in to biological tissues. Therefore, they have e merged as powerful diagnost ic medical imaging techniques for in vivo and non invasive detection of precancerous lesions Besides the advancement of photonics and fiber optics, system miniaturization, especially the miniaturization of the optical scanning mechanisms is the prerequisite for their development of cli nical endsocopic applications [13 15] While in the seeking of technical support, MicroElectro -Mechanical System (MEMS) a technology that enables the small devices and systems with the integration of electrical, mechanical and optical components and has significantly influenced many other areas, turns out to be the natural choice [16, 17] The primary objective of this project is to develop MEMS based transverse and translatory optical scanners on demand of the system miniaturization of the new biomedical imaging

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20 techniques, and thus the miniaturized imaging endoscopes for detecting and diagnosing in vivo precancerous lesions. This chapte r discusse s the new imaging technologies together with other potential application of the proposed MEMS o ptical scanners and then the research objectives and plan of this project 1.1 Conventional Cancer Detection Methods and Their Limitations A wide variety of cancer detection and medical test methods has been well -developed. The conventional non -invasive i maging techniques used for most cancer screening exam include ultrasound, magnetic resonance imaging (MRI), positron emission tomography (PET) scan, computer tomography (CT) scan, and some other techniques used for specific cancer type s such as mammography and colonoscopy [18] T he majority of cancers (>85%) are associated with morphological and functional alterations of cells inside epithelium layers of human body [19] t herefore, a fter the suspicious lesions being discovered and located, an incisional or excisional biopsy is usually needed for ex vivo histological and pathological s tudy of the tumor cells, to help determine the optimum treatment protocol for an individual patient However, the conventional imaging techniques are largely limited by their spatial resolutions, which are generally restricted to several millimeters range This makes them extremely difficult to detect and diagnose many precancerous lesions at curable stage, whose dimensions are even down to microns and far beyond the resolution capability of these conventional imaging modalities [18] Furthermore, for the inspection of internal organs, the currently used biopsy endoscopes are equipped with white -light cameras which can only visualize the surface morphological charac ter s but not detect the precancerous lesions under tissue surface. There are also some popular in vivo diagnosis techniques being investigated such as fine needle aspiration (FNA) and core needle biopsy (CNB), which have greatly facilitated the diagnosis process [20, 21] However, they are limited by the diagnosis ability, still invasive with

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21 discomfort and bleeding side -effects, expensive and time -consuming Most importantly, they share the common disadvantage of the conventional rand om biopsy lack of an accurate and high resolution imaging guide for biopsy sites identification. Therefore, for cancer detection in the early and curable stage, imaging technologies n millimeter ranges) are highly desirable. Also, the detection methods should be simple, efficient, inexpensive and noninvasiv e for increasing the extent of prevalence and frequency of cancer screening exams among large populations. 1 .2 New Biomedical Imaging Techniques 1.2.1 Optical C oherence Tomography Optical coherence tomography is a low coherence interferometry based non invasive imaging technology with micrometer resolution and cross -sectional imaging capabilities [3]. Sub -micrometer resolution has been achieved due to the introduction of wide bandwidth light source such as superluminescent diodes and femtosecond lasers [22] OCT obtains the subsurface information by imaging reflections from within tissue, thus is effectiv ely an optical ultrasound. It uses an optical interferometer with one optical beam delivered to a reference mirror and the other into the sample subject. The light penetrated into the subject will be scattered and absorbed; only the light reflected or sc attered from one depth is coherent with the reference light and thus can be detected by the interferometer and enables cross -sectional imag ing Figure1 1 shows a simplified schematic of a time-domain OCT setup. OCT attracts a great deal of interest in bio medical community mainly due to its crucial advantage of micrometer or even sub-micrometer high resolution ability with millimeter range imaging depth. This high resolution is far beyond conventional imaging modalities such as MRI or ultrasound. In additio n OCT is non invasive or minimally invasive and it can provide in vivo

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22 instant and direct imaging Furthermore no sample preparation is required enabl ing convenient, low -cost and safe use in office or clinic. Therefore OCT is very suitable to detect pr ecancerous lesions under epithelium layer at their very early stage, and has the potential to substitute the risky, costly and time -consuming biopsy procedures. Low coherence light source 50:50 Fiber coupler Optical delay line Photo detector Signal processing Computer & Display Transverse scanner Tissue sample Figure 1 1 S implified schematic of time -domain OCT setup One of the key components in the OC T system is the scanning mechanism in the sampling arm for transversely scanning the light beam over the sample surface to obtain the cross -sectional image ( 1 D scan) or a 3 D reconstruction of the tissue ( 2 D scan). S ome efforts on alternative imaging met hod of OCT ha ve also been investigated F or example, instead of transversely scanning a single point of the light, a new method known as full -field OCT (or parallel OCT) uses commercially available CCD camera to generate en face image over an illuminated area of the tissue [23, 24] However, the bulky microscope objective required in the sample arm limits the in vivo and endoscopic application of OCT. For the internal organs such as cardiovascular,

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23 gastrointestinal and pulmonary imaging application, miniaturized OC T endoscopes (in a few millimeters diameter) with scanning mechanisms and optical components including fiber optics and collimating lens must be developed. The scanning mechanism must be small enough to b e fitted into the endoscope, has decent scanning range and speed required by the imaging, and the multi -funct ionality of both 2 D transverse and axial scan is also desired. The detailed endoscope and scanner requirements will be discussed further in section 1.3. 1.2.2 Nonlinear Optical Microscopy Nonline ar optical microscopy (NLO or NLO M ) is another family of new rapidly -growing imaging technique s I t is based on the effects of higher order light -matter interactions, and thus fundamentally differs from the linear light -matter interact ion of traditional mi croscopy [6 9] The major nonlinear optical effects involved in NLO include multiphoton absorption (two-photon excitation fluoresecence TPEF) higher harmonic generation (second harmonic generationSHG and third harmonic generation THG) and coherent anti -Stoke Raman scatterin g (CARS) [25] All these nonlinear optical effects have demonstrated th e capability of high resolution 3 D visualization of tissue structure and thus have great potential for non invasive, in vivo biomedical imaging application s P articularly, TPEF and SHG are currently the most popular since femtosecond pulsed laser s in the near infrared wavelength range are used for excitation which can be greatly localized in time and space to maximize the fluorescence output [26 29] The major advantages of TPEF include 3 D localizatio n of the excitation volume thus providing inherently high resolution without the need of confocal detection optics; relative ly deep penetration depth of the near infrared light excitation source into turbid biological tissue ; and grea t spectral flexibility due to the simultaneous excitation of different fluorophores with a single wavelength. In addition, SHG as a coherent (phase preserving) process can produce highly polarized and predominantly forward -directed radiation instead of isotropic emission, which

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24 further enables the visualization of intrinsic structure s without using exogenous labeling and orienting the protein structures [14] Research efforts on the combination of TPEF an d SHG have stimulated new insights on cancer research [30 32] Ti: Sapphire Laser Prechirp Unit Dichroic Mirror 40X Objective lens PMT Iris Filter Bandpass Filter Double -clad Photonic crystal fiber Transverse scanner Tissue sample Ti: Sapphire Laser Prechirp Unit Dichroic Mirror 40X Objective lens PMT Iris Filter Bandpass Filter Double -clad Photonic crystal fiber Transverse scanner Tissue sample Figure 1 2 S implified schemat ic for TPEF and SHG imaging set up Figure1 2 illustrates a simplified nonlinear optical imaging setup that can be simultaneously use d for TPEF and SHG. T he re are two major challenge s for endoscopic imaging application of NLO The first is to efficiently deliver the excitation light as well as collect the fluorescence signal which relies in the advancement of fiber optics The second is still the miniaturization of the laser scanning mechanism for sample scanning and eventually the compact endoscope as in OCT. To obtain the real time 3 D visualization of tissue structure, the scanner must be capable of 2 D scan, and the aperture size has to be large enough ( ~1 mm) for efficiently collecting fluorescence signal but the device should be small enough (<2 mm) to be integrated in the endoscope. 1.2. 3 Confocal Laser Scanning M icroscopy Confocal laser scanning m icroscopy (CLSM) is an imaging tec hnique that has been widely used in numerous biological science disciplines and also clinically used in the evaluation

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25 of various eye diseases due to its ultra high resolution and the ability to collect serial optical sections from thick specimens [10 12] The principle of confocal microscopy is based on the optical rej ection of out of focus information by pinhole apertures [11] As illustrated in Figure1 3, the laser source passes through the light sour ce pinhole aperture and is focused into the specimen in a confined volumetric field; the returning fluorescent light from the specim en detection point is recollected by the lens and then reflected by the beam splitter, passes through the other pinhole aperture to the photo detector. The light source pinhole aperture can confine the detection point in the specimen to be less than one mi crometer in size, and the detector aperture suppresses the out of -focus light that is not coming from the focal point, thus images with much higher resolution than conventional fluorescence microscopy can be obtained. By displacing the focal plane in vario us axial positions below the tissue surface, the intact tissue can be optically sectioned and 3 D visualized based on this pixel by-pixel reconstruction. CLSM can provide a typical axial resolution of 3 5um, lateral resolution of 0.3 1um, and up to half m illimeter imaging depth, thus is promising for detection and clinical diagnosis of sub-cellular morphologic changes and functional alterations associated with the earliest cancerous stage [33, 34] In addition to the miniaturizatio n of optics, the major hurdle for the translation of CLSM to the in vivo endoscopic application is the scanning mechanism for fast image acquisition CLSM is based on the pixel -bypixel signal collection imaging method, which requires fast scanning mechani sm for scanning the light beam over the specimen volume. There are two commonly used methods for capturing image at high speed, the acousto -optic deflector (AOD) method [35] and the spinning Nipkow disk method [36, 37] .The first method has the disadvantage of wavelength dependence thus is limited for most fluorescence confocal imaging, while the bulky disk spinning mechanism of the latter method makes it not applicable for the in vivo endoscop ic

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26 application. Thus miniaturized beam scanning mechanisms must be developed Furthermore, besides the transverse scanning, for imaging depth in the specimen, a miniaturized axial scanning method to mechanically mov e the lens elements for displacing the focal plane into the specimen is also desired for integration and miniaturization of the endoscope system [15, 33] Fiber coupler Transverse scanner Tissue sample Primary image plane Scan lens Tube lens Objective lens Axial scan Laser PMT Fluorescence filter Pin hole Pin hole Figure 1 3 S implified schematic of a fluorescence confocal microscopy 1. 3 MEMS -based Endoscope Recently v arious miniaturiz ed scanning mechanisms have been explored following the rapidly -growing research popularity of endoscopic applications of the above mentioned new imaging technol ogies. The current scanning mechanisms can be divided into two groups, the proximal and the distal scanning based on the distance of the scanner to the light source. The proximal scanner is at the source end of the light delivery fiber and doesnt need to be capsulated into the small en doscope catheter T his type of scanners usually employs bulky galvanometer mirrors and requires fiber bundle technology [38 40] For example, Gobel demonstrated a two photon microscope with a 3 -mm diameter fiber bundle consisting of 100,000 individual fibers

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27 closely packaged, and a galvanometer mirror for raster scanning the proximal end of the bundle to translate the scan to the distal end then onto the specimen [40] However, the spacing and the thin cladding layer between the in dividual fibers limit the lateral resolution, cause leakage of excitation beam and reduce the i maging contrast [14] The distal scanning is usually realized by scanning mechanisms inside the endoscope to either scan the fiber tip or t he light coupled from the fiber for exam ple, rotating the entire cable that carries the optical fiber [41] or using piezoelectric or electromagnetic actuators to swing the distal fiber tip [42 46] In addition, to image some hollow internal organs, circumferential scanning endoscopes employing rotating micromotors have also been developed [47 50] The other distal scanning method is based on miniature scanners built by MEMS technology [51 63] MEMS is a new t echnology developed from the microfabrication processes of conventional Integrated Circuits (IC) technology enabl ing the fabrication and system integration of actuators and sensors in micrometer scale. Micromirrors based on this technology are provided with the merits of miniature size, low power, fast speed and potentially low cost and thus they are very suitable for endoscopic imaging applications. Electrostatic micromirrors developed by several research groups have demonstrated applica tions of in vivo endoscopic imaging in OCT [56 58] N LO [59 ] and CLSM [60] Electrostatic actuation can typically provide operating frequency above 10 k Hz and is preferable for fast imaging acquisition. However, it usually requires high driving voltages from tens of volts up to hundreds of volts and thus prompts safety issue for in vivo internal organ imaging Moreover both the deflection angle s in the angular scanning mode and the vertical displacement s in the piston mode are small, only up to about 10 degree and tens of mi crometers respectively [55 63] which further limits electrostatic actuation to be impractical in large motion applications.

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28 The other popular actuation method, el ectothermal actuation usually provides large actuation range s (tens of degree rotation angles and hundreds of micrometers piston displacement) at much low er driving voltages (less than 20 V) and acceptable operating frequenc ies up to a few k Hz Xie et al reported a 5 -mm diameter OCT endoscope using a 1 D electrothermal MEMS mirro r where high resolution cross -sectional OCT images were demonstrated in 2002 [64] 2 D electrothermal mirrors developed by Jain et al. was then successfully used in a bench top nonlinear optical imaging system to obtain 3D visualization of intact tissue [53] A n endoscope prototype integrated with an electrothermal large vertica l displacement lens scanner as a focusing mechanism, for optical coherence microscopy and confocal microscopy has also been demonstrated [54] The re are still some drawback s in the above mentioned electrothermal micro scanner designs incl uding the significant lateral shifts of the aperture center both in scanning mode and piston mode of large vertical displacement scanners, the low fill -factor, i.e. relatively small effective aperture size compared to the devi ce footprint Furthermore, a tip tilt -piston (TTP) micromirrors with centroidal rotation motion is desired in compact en doscope for larger effective aperture size and efficient beam collection, a nd new designs capable of f ull -circumferential scanning (FCS) is highly desired for endoscopic imaging of hollow internal organs. 1. 4 Research Goal and Tasks The main research goa l of this dissertation is to develop electrothermal based MEMS optical scanners and miniaturized imaging endoscopes with the optical scanners integrated for in vivo detection and diagnosis of precancerous lesions. To reach this goal t he following research tasks will be completed First to develop a new el ectrothermal actuator capable of large vertical displacement with zero lateral shift and tilting, namely lateral -shift -free large-vertical -displacement (LSF -LVD)

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29 electrothermal actuators and to e xtend t he vertical scanning range of the new actuator to be above 1 millimeter range This task is an improvement of the LVD electrothermal actuator developed by Jain et al. [65] B oth LSF LVD micromirror an d microlens scanners will be fabricated. O ther potential application s of LSF LVD micromirrors such as miniature Fourier transform infrared spectrum (FTIR) analyzer will be investigated Secondly, to design and fabricate TTP micromirrors with higher fill fa ctor s and a centroidal rotation mode based on the new LSF -LVD microactuator ; and to develop OCT and NLO imaging endoscopes for in vivo biomedical imaging applications. The scanning mirror should be capable of tip -tilt optical angles above 30 and piston s canning range above half millimeter and driving voltages less than 10 V. The aperture size should be at least 1 mm by 1 mm with a device footprint less than 2 mm by 2 mm to be fitted into the miniature endoscope with a diameter less than 5 mm. Last ly to d evelop a new micromirror for a new full -cir cumferential scanning endoscope design. The new FCS micromirror should be able to mechanically rotate 45 at real time imaging speed and is dual reflective with each mirror side enabling half circumferential scan range. A novel actuator will be designed to achieve the large scan range, and a new self aligned deep trench etching process will be developed to fabricate the dual -reflective micromirror A similar fabrication process will be investigated to develop ligh t -weight structure of the single crystal silicon (SCS) mirror plate for micromirrors with large aperture size (>3 mm) 1. 5 Dissertation Overview Th ere are six chapters comprised in this dissertation The first chapter introduces the background and motivat ion of this project, including the new biomedical imaging technologies

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30 for early cancer detections, the corresponding miniature endoscopes currently developed and the proposed research objectives and task s In Chapter 2, a comprehensive literature review of currently developed MEMS micromirrors is firstly presented, and then the design, principle of operation and modeling methods of electrothermal actuation is discussed in detail, followed by the review of the prior work on electrothermal micromirror desi gn and a summary of remaining problems. Chapter 3 presents a novel LSF -LVD electrothermal actuator design which is capable of generating lateral -shift and tilting free large ve rtical displacement s and the LSF -LVD actuator based micromirror s and microlen s scanner s. The actuator design issues are analyzed by lumped element models and a n improved LSF -LVD design with faster scanning speed is also presented. In Chapter 4, a novel full -circumferential scanning endoscope design and a dual -reflective micromirror based on a novel dual -folded bimorph (DFB) design for the f ull -circumferential scanning are presented A nove l fabrication process for the dual reflective mirror and light weight SCS micromirrors is also presented. Chapter 5 presents the applications of the developed optical scanners First, the imaging endoscope development including 2 D scanning endoscopes using TTP scanning micromirror s for OCT and NLO imaging, and a microlens scanner based on a modified, more robust LSF LVD actuator design capable of scanning glass lens es are presented Other potential applications of the developed MEMS optical scanners are also investigated, such as, TTP mirror arrays are demonstrated for optical phased array applications, and large aperture mirrors by cascaded LSF LVD actuators for free -space optical communications is also presented. Lastly as another important application of the LSF LVD micro mirrors, miniature Fourier transform spectrometers based on translatory scanning mirrors by modified LSF LVD actuator is dem onstrated

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31 Finally the entire research efforts and a list of accomplishments will be summarized in Chapter 6, together with suggestions for future work and potential research opportunities of the project.

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32 CHAPTER 2 MEMS ELECTROTHERMAL OPTICAL SCANNERS As already discussed in the first chapter, for the clinical use of OCT, N LO and CLSM imaging techniques in early cancer detections and diagnosis, system miniaturization of the imaging endoscopes is one of the key obstacles The rapidly developing MEMS technology as a great breakthrough enabling the micro -scale devices and systems has shown its great impact in almost all the disciplines [16, 17] and thus is naturally the first choice for realiz ing miniature optical scanning mechanisms. This chapter first reviews the current MEMS optical scanners based on dif ferent actuation methods and introduces the preferable selection of electrothermal actuation for in vivo biomedical imaging applications. It then presents the design, principle of the electrothermal bimorph actuation followed by an introduc tion of the prior actuator designs and a summary of their remaining problems 2 1 MEMS Optical Scanners MEMS optical scanners can be divided into two categories, reflective and refractive scanners. MEMS reflective optical scanners, also referred as MEMS scanning micromir rors have been widely investigated in a variety of applications such as optical communication [66] optical switching [67 70] optical display [71, 72] bea m steering [73, 74] and endoscopic imaging [51 63] Fo r biomedical imaging applications the high spatial resolution requires a flat mirror surface (radius curvature > 0.3m) and a large aperture size (>0.5mm). Micromirrors based on stressed thin film reflectors as some commercially available products, such as Texas Instruments DMDs (Digital Micromirror Devices) [75] and Lucent Technologies optical switch [67] are not practical for biomedical imaging due to their small aperture size s (~0.1 mm). S ingle crystal silicon (SCS) micromirrors based on bulk -micromachining process es provide improved mirror flatness with relatively large aperture size s and thus have been preferably developed via

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33 various actuation s including electrostatic, electrom agnetic piezoelectric and electrothermal methods. The early research on electrostatic-driven SCS micromirrors can be traced to almost three decades ago when Peterson et al. demonstrated the parallel -plate electrostatic tor sional SCS micromirror in the year of 1980, wit h a maximum rotational angle of 2 at resonance and 300 V driving voltage [76] It utilizes the ele ctrostatic force between two electrodes which are locate d under the susp ended mirror plate and on the bottom parallel plate respectively to generate the rotation More recent ly developed micromirrors using this approach have yielded larger rotational angels up to 8 with smaller driving voltages ranging from 40 V to 200 V [67, 7781] However, the achievable rotational angles of the parallel -plate method are greatly limited due to the inherent trade-offs between the mirror siz e, gap distance and driving voltage. Another new design employing vertical comb drives (VCD) was then developed to improve the rotational angles and mirror aperture size as well [82 85] The most common structure used is a gimbaled structure in which torsional beams are used to suspend the mirror plate and transform the rotation generated by vertical comb drives. For example, a resonantly excited 2 D micromirror based on SOI wafers and Si wet etch bulk micromachining process was demonstrated by Schenk et al. in 2000, w here rotational angles of 5.5 were obtained with resonance excitation at 16 V [82] Conant et al. developed staggered comb drivers for a 1 D mirror through wafer bonding and achieved 6.25 rotational an gle at resonance [83] To further improve the actuation range of the comb drives, a ngular comb drive design s have been demonstrated [86 88] For example Patterson et al. used photoresist reflow to form the hinge and an initial tilted a ngle of the moving combs [86] and later photoresist was replaced by polysilicon latches [87] with rotational angles of 2 at 120 V and 6.2 at 55 V respectively Xie et al. also demonstrated a curled -hinge comb

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34 drives utilizing the initial curling of the stressed thin films and achieved 4.7 at 20 V [88] These comb -drive gimbaled designs allow larger mirror size and comparable actuation range with the parallel plate designs. However, the rotational angles are still less than 10 due to the comb drive s stroke limited rotation Also for the electrically isolated mech anical decoupling in 2 D micromirrors the trench filling and polishing process complicated the fabrication [89] A new laterally actuated gimba l less structure was demonstrated by Milanovic at al. with the angular rotation being mag nified to >12 at resonance with < 150 V driving voltage [90] Th e same design can achieve a piston motion as well. More recently, Tsia et al. reported a radial vertical comb driver actuator design to elimina te the need of gimbals and obtained 5.6 rotational angle at 61 V [91] A pre -stressed comb drive design by Chiou et al. also demonstrated rotational angle up to 6.5 and piston displacement of 45 [92] High resonant frequencies of these electrostatic actuators make them ideal in the applications where fast scanni ng is desired. However, the small scanning range and high driving voltages that will prompt safety concern s still confine their applications for in vivo endoscopic imaging. Also, most electrostatic mirrors have small fill factor s due to the large area take n by the electrostatic actuators and the mirror aperture size is usually limited to 0.5 mm. E lectromagnetic micromirrors uses the Lorenz force for mirror actuation which has the advantages of relatively large actuation force s (on the order of millinewton s) at low driving voltages and long effective range (up to millimeters) Also electromagnetic force can be either attractive or repulsive [93] Currently developed electromagnetic micromirrors usually employ an external magnetic field with a magnetic material such as permalloy [93 96] or electrical coil s with flowing current [93, 94, 97, 98] integrated in the mirror plate for actuation. Although large r rotational angles up to 8 0 were o btained for1 D micromirrors [96] and some 2 D micromirrors

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35 demonstrated 15.7 [97] 18 [98] and 22 [95] actuation range s the bulky components required for external magnetic field stringent ly re strict the device miniaturization and the packag ing compact ness required for endoscopic imaging applications. On -chip coil design to eliminate the need of external magnetic field was also reported [94] but it is difficult to realize 2 D scan th rough this method and it involves high power consumption for generating the driving magnetic field. Recently p iezoelectric actu at ed micr omirrors have also attract ed some research interests especially after thin film deposition technology of piezoelectric PZT (lead zirconium titanate ) films rea ched an advanced state [99] The piezoelectric actuation is based on the bending of the metal/PZT/metal sandwiched unimorph beam [100103] or double layered PZT [104] corresponding to applied electric voltages across the PZT material. Considera ble tilting angle s have been demonstrated with better linearity and much lower driving voltages than electrostatic micromirrors, and high bandwidth is also achievable due to the fast response [105] However, the scanning ranges of c urrently developed piezoelectric micromirrors are still limited, with typical rotational angles ranging from 2 to13 [100102, 104] and piston displacement only a few micrometers [106] Also, charge leakage and hysteresis problems are also the remaining obstacles for practical application of this actuation method. Elect rothermal bimorph actuation provides great e r actuation force and larger deflection range than similarly -sized actuators by other method s such as electrostatic and piezoelectric actuation. Other advantages include almost linear deflection versus power relat ionship, relatively simple fabrication process es and structure designs that are easy for system integration. A bimorph beam consist s of two materials with different coefficient of thermal expansion s (CTE ), T he actuation principle is based on the bending motion of bimorph beams in response of

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36 temperature change that is usually introduced by Joule heating. Researches on bimorph actuators based on different materials have been reported [107114] Buser et al. reported a biaxial scanning mirror that uses Al uminum (Al) and Si licon (Si) for bimorph materials and doped Si as the guided electrical current path for Joule heating [107] Bhler et al. demonstrated Al/ SiO2 bimorph beams based on a CMOS fabricati on process [108] Other materials such as metal oxide [109] and polyimide [110, 111] were also used for bimorph actuation. Large rotation angles up to 40 for both 1 D and 2 D scanning mirrors at driving voltages less than 20 V have been demonstrated by Jain et al. [52, 115] T hese devices are fabricated by a post -CMOS MEMS proc ess and use Al/SiO2 bimorph beams with embedded P oly -Si as the heating material. Actuators using thermal expansion of single material such as Si were also reported [116, 117] The main drawbacks in electrothermal actuation include relatively high power consumption and low bandwidth due to the slow thermal response and poor temperature stability However, for applications th at require large scan range with low driving voltages such as endoscopic imaging, electrothermal actuation turns out to be a very sui table approach. The simple structure allows large aperture size on small device footprint and ease of system integration. Also, the power consumption is sufficiently low compared to conventional macro-device s and, as shown later the device s can be operat ed at reasonable high resonant frequencies ranging from hundreds of Hz to a few k Hz. As mentioned in section1.2.3, in addition to reflective scanning mirrors r efractive optical scanner s are also desired for the axial focusing mechanism of endoscopic imag ing systems. The depth scanning in CLSM and optical coherence microscopy (OCM) usually requires millimeter range axial displacement of the imaging focal plane. There are currently two approaches that can displace the focal plane of the focusing lens. One m ethod is to use tunable lenses of which the

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37 focal length s can be varied axially [118126] The reported tunable lenses are mostly liquid or polymer lenses with deformable shape based on electrowetting principle [121, 122] pneumatic actuation [123] microfluidic actuation [119, 126], or a new hydrogel material [124, 125] But these tunable lenses are not practical to use for endoscopic imaging due to the required large actuation voltages slow response, or the bulky sizes. The other method u sed for the focusing mechanism is to mechanically displace the focusing lens along the optical axis by microactuators to obtain the axial scan of the focal plane inside the tissue [15, 33, 127] Microactuators capable of vertical displacement based on electrostatic [61, 128] pneumatic [129, 130] and electrocapillary [131] actuation methods have been reported but the ir achievable scanning range s are too small to meet the requirement for in -depth imaging applications. The electrothermal actuation on the other hand is very promising for generating large vertical displacement. For example, e lectrothermal LVD actuators developed by Ja in et al. already demonstrated over half millimeter vertical displacement range [132] As a co nclusion of the above reviews, electrothermal actuation meets all the requirements for endoscopic imaging applications, including large actuation range s low driving voltages, relatively large aperture size with small device footprint, and reasonable scann ing speed. The principle of electrothermal bimorph actuation will be presented in the next section, followed by an introduc tion of the prior designs and their remaining problems. 2 2 Principle of Electrothermal Bimorph Actuation The electrothermal optica l scanners in this proposal are all based on electrothermal bimorph actuation. The bimorph structure is a stack of two layers of different materials with di fferent CTE s The cantilevered bimorph beam can co n vert the small strain difference between the two layers to a large displacement perpendicular to the strain plane. The actuation principle of electrothermal bimorph beams can be simply explained as following: with the temperature

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38 change induced by Joule heating of an electrical current, the difference of thermal stresses between the two layers due to different CTE s leads to the bending of the cantilevered structure 2. 2 .1 Stress and Curvature in Bimorph Cantilever The stresses in the thin films include extrinsic stress and residual internal stress. The extrinsic stress is from the external factors such as packing, and the thermal stress due to the different CTE between the thin film and substrate or other thin film material. There are two types of residual internal stress es compressive and tensile stre ss. The compressive stress leads the thin film to have an in plane expansion tendency, while the tensile stress tends to contract the thin film parallel to the substrate. The residual stress mainly consists of the thermal stress, resulting from deposition temperature conditions and the intrinsic stress formed during the film nucleation [133, 134] Through the measurement of surface curvature by a wafer curvature (disk) method, t he stress of thin films can be derived from the Stoney equation [135] as : ) 1 1 ( ) 1 ( 62 s sf f s s s fr r t t E (2 1) sE and s are the Youngs Modulus and Poissons ratio of the substrate material, st and ft are the thickness of the substrate and thin film respectively and sr and sfr are the radius of curvature before and after the thin film deposition, respectively. Bimorph canti lever beams usually ha ve an initial curling after being released from the substrate. This curling is beneficial for generating a large out of -pla ne initial rest position to achieve large scan range for optical scanners Due to the fact that the tensile stressed material tends to contract while the compressive stressed one tends to elongate, the bimorph beam with the tensi le material on top curl s upward while the bimorph beam with the compress ive material

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39 on top curl s downward, as illustrated in Figure 2 1 This initial curvature or as related to the residual stresses can be expressed by [136] : 1 1 3 2 2 2 2 3 1 1 2 1 2 2 2 1 2 1 2 16 4 4 ) ( )( 6 1 t E t E t E t E t t t t t t ro o o (2 2 ) 1t and 2t are the thickness of two layer s, respectively 1 o and2 o are the linear stains due to their residual stress es and 1 'E and 2 'E are the b iaxial elastic modulus which is related to Youngs modulus and Poisson ratio: ) 1 /(' E E (2 3) Let denotes: 1 1 3 2 2 2 2 3 1 1 2 1 2 2 2 1 2 2 16 4 4 ) ( 6 t E t E t E t E t t t t t t (2 4) and o as the difference of linear strain due to residual stress t he initial radius can be simplified as: o b ot r 1 (2 5) The above analysis applies to thermally induced curvature as well. The thermally induced strain for each material can be expressed: Tt (2 6) is the temperature coefficient of expansion. The different strain resulting from CTE difference leads to the thermally induced curvature: T t t rb t b t 1 (2 7)

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40 The initial curvature and the thermally induced curvature are additive and result in the final bending r adius: Figure 2 1 A side view of the bimorph beams with initial upward curling t or r r 1 1 1 (2 8) Thus the final tilting angle at the bimorph tip which is approximate to the arc angle of the curved beam can be expressed as Equation 2 9 Note that thermally induced curvature is negative if the tensile stressed material has larger CTE than the compressive one because the heated bimorph will unfold and bend backward to the substrate. t o o b b bT L t r L ) ( (2 9) o is the initial tilting angle and t is the actuation angle by temperature change. The temperature change along the bimorph beam is not ideally uniform thus the actuation angle is usually expressed by the average temperature change on the bimorph beam [137]

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41 T L tb b t (2 10) A n equivalent force generated at the tip of the bimorph beam that is induced by the temperature change can also be defined [138] The deflection of the bimorph beam due to temperature change can be counteracted by an assumed external force eqF applied at the tip to keep the beam at a fixed position. This external force can be related to the tip deflection b of the cantilever beam by: b b eqK F (2 11) bK is the spring constant of the cantilever beam, as being well known, is given by: 3 34 l Ewt Kb (2 12) E is the equivalent Youngs Modulus of the bimorph beam, w, t and L are the beam width, thickness and the length. As for small deflection, it can be approximately related to the thermally induced bending radius tr and eventually to the temperature change by plugging in the thermally induced curvature (Equation 2 7 ) as: T t l r lb t b 2 22 2 (2 13) This yie lds the external force ( Equation 2 11) to be: T l t w E Feq 82 (2 14) 2. 2 .2 Bimorph S ensitivity and M aterial S election From the linear relationsh ip between temperature change and actuation angle as shown in Equation 2 10, the bimorph actuation sensitivity can be simpl y defined as the achievable actuation angle per degree Kelvin:

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42 b b tt L T S / (2 1 5 ) which is linearly dependent on the bimorph beam length and CTE difference of the bimorph materials. Other affecting parameters include beam thickness of each layer and mechanical properties in volved in the curvature coefficient As expressed by Equation 2 4 the curvature coefficient is strongly dependent on the thickness ratio for given materials. By defining thickness ratio 2 1/ t t and the ratio of the biaxial elastic modulus of each layer 2 1/ E E the unit less curvature coefficient can be expressed as: ) 2 3 2 ( 2 1 ) 1 ( 62 3 2 (2 1 6 ) F or a given total thickness, the condition to meet the maximum value of curvature coefficient can be easily fou nd by setting the partial derivative of with respect to the biaxial modulus ratio equal to zero [136] as to be : 12 or 1 2 2 1E E t t (2 1 7 ) Under this condition, the curvature coefficient becomes a constant value of 1.5 for all material This simplifies the material selection and steady -state mechanical optimization, the materials should be selected based on a high CTE difference. Table 2 1 summaries the common ly used bimorph material with their properties. As can be seen from Table 2 1, Al and SiO2 has the relatively large CTE difference thus can provide high actuation sensitivit y. Both Al and SiO2 are among the most common materials used in MEMS micromachining and are cost efficient. Also, Al is a good reflective coating material for a wide band of light sources Therefore, Al and SiO2 are selected for the bimorph actuator materi al in

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43 this proposal. As for the heater, Poly Si is a convenient material. However, the embedded polysilicon heater exhibits hysteresis problems and self annealing effect s which limit their stable scanning ranges. For instance, a hystersis problem was obser ved in [64] and self annealing effects of Poly -Si embedded bimorph beams causes an unsta ble scan at high voltages in [52] Instead of polysilicon, platinum (Pt) is another good choice that has been widely used in MEMS gas sensors and anemometers due to its long t erm stability of mechanical and electrical properties at high temperatures [139, 140] it also has relative ly large t emperature coefficient of resistivity (TCR) compared to polysilicon thus has potentially better sensitivity to be used as a temperature monitor on electrothermal actuators Table 2 1 Thermomechanical properties (at room temperature) of materials possible for bimorph actuation Material Coefficient of Thermal Expansion 106/K Youngs Modulus GPa Thermal Conductivity W/mK Specific Heat 103J/KgK Density 103 Kg/m Si 2.6 162 170 0.691 2.42 SiO 2 0.4 740 1.1 0.84 2.66 Si 3 N 4 2.8 155 18.5 0.711 3.19 SiC 3.5 457 86.5 3.2 Poly Si 2.3 160 0.754 2.33 Al 23.0 690 235 0.9 2.692 Au 14.3 800 318 0.129 19.4 Cr 5 14 0 90.3 0.447 7.19 Cu 16.7 120 401 0.387 8.95 Ni 12.8 210 91 0.444 9.04 Pt 8..9 147 73 0.133 21.5 Pb 28.7 160 35 0.128 11.48 Ti 8.6 116 21.9 0.523 4.51 Material properties obtained from [116] and material database of memsnet.org 2. 2 3 Electrothermal A nalysis An electrothermal model for a 1 D micromirror design has been developed by Tod d and Xie [137] The 1 D electrothermal mic r omirror is illustrated in Figure 2 2, where the Al c oated

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44 mirror plate is connected by a series of Al/ SiO2 bimorph beams anchored to the substrate at the other end. The bimorph beams ha ve a thin layer of Poly-Si in between for Joule heating. As the heat transfer on the bimorph beams in y and z direction is only convection though air if the radiation is neglected, the problem can be simplified to 1 D and the temperature is only a function of x along the bimorph beam. The heat transfer equation on the bimorph beams can be expressed as: Substrate SCS mirror plate Al SiO2Poly -Si Si Substrate thermal isolation Mirror thermal isolation T0 T0 RTLRTRLb Substrate SCS mirror plate Al SiO2Poly -Si Si Substrate thermal isolation Mirror thermal isolation T0 T0 T0 RTLRTRLb Figure 2 2 Schematic of 1 -D micromirror design illustrating electrothermal model t t x T k t x g t x T t k h x t x Tb b b ) ( 1 ) ( ) ( 2 ) (2 2 (2 1 8 ) with the steady -state equation reduced to: 0 ) ( ) ( 2 ) ( b b bk x g x T t k h x T (2 1 9 ) bk and are the thermal conductivity and thermal diffusivity of the bimorph beam, h is the convection coefficient and g denotes the heat generation, i.e., the electric power generated by Joule heating per unit volume expressed as: )] ( 1 [ ) (0 2t x T J t x gR E (2 20)

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45 in w hich J is the electric current density through the heater, 0 E and R are the electrical resistivity of the heater at the initial temperature and its temperature coefficient of resistivity T he boundary conditions applied for solving the differential equation s i nvolve the equivalent thermal resistances seen by the bimorph beams at both ends denoted by TLR and TRR and the thermal resistance of the bimorph b eam, TbR i.e., Tb TL bR R L x t x T t x T ) 0 ( ) 0 ( (2 2 1 ) Tb TR b b bR R L x t L x T t L x T ) ( ) ( (2 2 2 ) A steady -state lumped element model based on the determination of the maximum temperature point on the bimorph beams by an analytic analysis was also developed to simplify the derivation of average temperature on bimorph beams [137] 2 3 Prior Electrothermal Actuator Designs Various MEMS optical scanners based on electrothermal bimorph actuatio n have been successfully developed by Xie and Jain [52, 65, 115, 132, 141] and demonstrated great potential for endoscopic imaging applications [51, 53, 54] These devices are fabricated by a DRIE Post CMOS MEMS process, using Al and SiO2 as the bimorph ma terial s and PolySi for Joule heating. To maintain the flatness of the micromirror with relatively large aperture size, single crystal silicon (SCS) is used to form the mirror plate. As shown in Figure 2 3 A ), the 1 D electrothermal micromirror design cons ists of a series of Al/ SiO2 bimorph beams and a bulk Si supported mirror plate connected to the tip of the beams. The small width of the thin film beams allows fast and complete undercut of Si underneath in an isotropic Si dry etch process and leave s the m ajority of bulk Si under the mirror plate. The mirror plate is coated by Al for high reflectivity. The Al on -top bimorph beams has an

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46 upward initial curling due to the residual stress and unfolds downward once being heated by applying a voltage to the embedded Poly -Si heater, and thus a rotation motion of the mirror plate around the bimorph beams then the scanning of the mirror surface can be obtained. A 2 D micromirror design as the extension of this 1 D design is illustrated in Figure 2 3 B), in which two sets of bimorph beams are orthogonally oriented and connected by a bulk Si supported frame with the mirror plate connected inside the frame. The two sets of bimorph beams are actuated by separate Poly Si heaters so two perpendicular axes of mirror rotati on can be achieved. The fabricated devices demonstrated static rotation angles of 31 at 18 V or 9 mA for the 1 D micromirror and 40 at 15 V, 25 at 17 V for the 2 D scanning micromirror [52, 141] Figure 2 3 E lectrothermal micromirror design and SEM pictures of fabricated devices A )1 D and B) 2 D [52, 141] Although large rotation angles can be obtained, the above micromirror designs still have the drawbacks such as the unidirectional scan, non -stationary center of rotation and large initial tilt angle that complicates the device packagin g and optical design of the imaging endoscope. A new large -vertical -displacement (LVD) electrothermal actuator design was then proposed to resolve the above issues [65] Figure 2 4 A ) shows the schema tic of the LVD actuator design I t A B MirrorSubstrateBimorph actuator Embedded Poly Si heater MirrorSubstrateBimorph actuator Embedded Poly Si heater MirrorSubstrateFirst axis Frame actuator Second axis Mirror actuator Frame MirrorSubstrateFirst axis Frame actuator Second axis Mirror actuator Frame

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47 uses two sets of complementarily -oriented bimorph beams in a folded structure that allows the mirror plate to remain parallel to the substrate after being released, with an initial elevation resulting from the large strok e lengths of the actuators. The same design can also achieve bidirectional 1 D scan with the two sets of bimorph beams being actuated independently. A bidirectional rotation angles from 16.5 to 26.5 at less than 5.5 V Based on this LVD actuator design, a lens scanner with a polymer micolens integrated was also developed and up to 7 1 large vertical displacement was achieved. As an extension of this LVD act uator concept, a 2 D LVD micromirror design was also developed. As shown in Figure 2 4 B), by replacing the mirror plate in Figure 2 4 A ) with another set of LVD actuator in an orthogonal orientation to the first set, the scan of the mirror plate can be actuated in both perpendicular directions, and a piston motion can be obtained as well by properly actuating the four bimorphs simultaneously. Optical scan angles of 40 and 30 in both directions are obtained at less than 12 V [65, 115] Figure 2 4 LVD electrothermal micromirror design and SEM pictur es of fabricated devices A) 1 D and B) 2 D [65, 115] A B Substrate Mirror Substrate Mirror SubstrateFrame actuator FrameMirror actuator Mirror SubstrateFrame actuator FrameMirror actuator Mirror

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48 2 4 Summary of Remain ing Problems The prior bimorph actuator designs as discussed in the previous section have achieved both large scanning angles and large vertical displacement as required for the most endoscopic imaging techniques. But there still remain some problems in th ese designs Substrate FrameMirror plate Frame actuator Mirror actuator Lf Ls Z Lf Substrate FrameMirror plate Frame actuator Mirror actuator Lf Ls Z Lf Figure 2 5 Cross -sectional view of LVD actuator showing the lateral shift Although the LVD actuator designs solve the unidirectional scan and large initial tilt problem s in those preliminary 1 D and 2 -D micromirror designs they bring in a new lateral shift problem As shown in Figure 2 5 in the piston motion of the micromirror, the mirror center shifts laterally as well. If the bimorph beams are actuated from the initial tilt position back to flat and parallel to the substrate surface, i .e., the actuation angle is equal to the intial tilt angle, th e vertical displacement Z and the lateral shift of the mirror center Ls related to the effective frame length Lf can be simply expressed as : sin fL Z (2 2 3 ) ) cos 1 ( f sL L (2 2 4 )

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49 Thus the lat eral shift is proportional to the actuated vertical displacement for a given actuation angle of the bimorph beam. For example, the LVD microlens in [142] can generate a maximum 0.71mm vertical displacement, but the lateral shift is as large as 0.42mm, which will greatly distort the microscopic image The lateral shift in LVD micromirrors also significantly reduces the effective optical aperture size. Another disa dvantage of the previous LVD designs is that a certain ratio of the driving voltages for the two complementary actuators needs to be maintained in order to obtain a n equal but opposite rotation angles of the two actuators thus a pure piston motion of the m irror plate without tilting and this experimentally -determined ratio is not constant but increases with the applied voltages [65] This greatly complicates the driving of the device especially for th e dynamic vertical scan Also the fill-factor of the LVD micromirror is relatively low due to the large area taken by the frames and the bimorph beams. Especially in the 2 D LVD micromirror design, the effective aperture size is only 0.5 mm by 0.5 mm for a given device footprint of 2 7 mm by 1. 9 mm with a fill -factor (the effective mirror area to the total area with its surrounding actuators) of only about 4.8 % This limits the further device miniaturization without sacrificing the aperture size of the mi rror In addition, the problem of non -stationary mirror center during actuation remains un solved W ith the hinged rotation motion of the mirror plate the lateral shift of the mirror center still exists and can easily cause the optical m isalignment in mini ature endoscopes thus further decreases the effective aperture size. A novel actuator design as a solution of all these remaining problems in current LVD actuators will be presented in Chapter 3. Furthermore, as mentioned in C hapter 1 a f ull -circumferent ial scanning (FCS) mechanism is highly desired for fast and efficiently imaging some hollow internal organs However, in

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50 current devices there is still a lack of a micromirror design capable of f ull -circumferential scanning. Also, the prior micromirrors ar e all based on the SCS supported structure to maintain the flatness of the mirror surface. But the mass of the bulk silicon will significantly increase with the mirror size, which may engender significant decay of the devices agility or may even surpass t he actuation force limit of the bimorph beams A fabrication method to make flat and large aperture mirror s but with light -weight structure s needs to be investigated. Chapter 4 will be focusing on the FCS micromirror and light -weight large aperture micromirror design. Lastly, the previous devices are all fabricated by DRIE Post CMOS process at die level. For the eventual commercial volume mass production, the feasibility of a wafer level fabrication process must be investigated. The nonCMOS fabrication pr ocess also eli minates the constraints on material selection and structure parameters by CMOS foundry

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51 CHAPTER 3 LATERAL SHIFT FREE LARE -VERTICAL DISPLACEMENT ELECTRO THERMAL ACTUATOR AND MICROSC ANNERS A s a solution to the remaining problems of the electro thermal LVD actuator described in Chapter 2 a novel LVD electrothermal actuator design with negligible lateral shift and tilting will be presented in this chapter. The innovation of this novel lateral -shift -free (LSF) LVD electrothermal actuator design is to use series -connected rigid frames and curled beams for self compensation of the lateral shift and tilt from each component but still take the advantage of large vertical displacement at the tip of the tilted frames, resulting in a lateral -shift -free an d tilt free piston motion at the tip of the actuator. By employing multiple LSF LVD actuators symmetrically oriented at the four edges of a mirror plate, this novel actuator design can achieve centroidal tip tilt scanning motion T herefore it is also prefe rable as the transverse scanning mechanism for endoscopic imaging. Various vertical actuators based on different actuation mechanisms have been developed [61, 92, 130, 143148] For example, Kwon et al. used an electrostatic vertical c omb drive design to generate a maximum 55 [61] Chiou et al. demonstrated an electrostatic micromirror capable of a maximum 45 m vertical displacement [92] Yeh et al. presented a thermal actuator based on a single layer step bridge that can move upward by about 13 [147] These above -mentioned actuators are limited in the endoscopic imaging applications due to their small actuation range (tens of microns) and high dr iving voltages (usually tens or even hundreds of volts). Some actuator designs with larger vertical displacement s have also been reported. For instance, Werber et al. demonstrated a thermo -pneumatic microactuator, with a maximum vertical displacement of ab out 385 [130, 148] However, this thermo -pneumatic device is very slow (time constant of about 6 seconds in [148] ), and its configuration is complex.

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52 In this chapter, the LSF LVD electrothermal actuator design is first presented, followed by demonstrating various fabricated devices including TTP micromirrors and LVD micro lens scanners And then a transient thermal analysis based on a lumped element model (LEM) and an improved actuator design with faster thermal response is also demonstrated. 3 1 LSF -LVD Electrothermal Actuator Design The three -dimensional ( 3 D ) model of the LSF actuator design is shown in Figure 3 1 The actuator is composed of three sets of Al / SiO2 bimorph beams with two frames connected in between. A platform (a mirror plate or a lens holder) is connected to the last set of the bimorph beams. The frames and the platf orm have single -crystal silicon underneath to provide structural rigidity and flatness P t heaters are embedded in all the bimorph beams for a uniform and efficient heating. SiO2 meshes are used to provide good thermal isolation. The bimorph beams curl up aft er being released to result in an initially -elevated platform, and will bend down ward with all three bimorph actuators being actuated By using three bimorph actuators and two frames connected in series and properly choosing their lengths, a purely vertica l piston motion of the platform with nearly zero lateral shift and tilting can be obtained. Substrate Bimorph I Frame II Bimorph III Frame I Bimorph II Platform 2 1 3 Substrate Bimorph I Frame II Bimorph III Frame I Bimorph II Platform 2 1 3 Figure 3 1 3 D model of LSF LVD actuator design with nomenclatures

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53 Figure 3 2 Cross -sectional view s chematics A) t wo -bimorph LVD a ctuator design with large lateral shift and B) t hree -bimorph LVD actuator design with no lateral shift Fig ure 3 2 shows the comparison between the two -bimorph LVD actuator design and the three bimorph LSF -LVD actuator design. In the two -bimorph LVD actuat or design (Figure 3 2 A )), the large vertical displacement is achieved by the angular amplification provided by the long arm length of the frame. The tilting of the platform can be compensated by the equal angular rotations of the two bimorph bea ms in the opposite directions. However, the rotation of the single frame will cause a significant lateral shift on the platform center, which is proportional to the vertical displacement as discussed in section 2 4. The new LSF -LVD actuator design (Figure 3 2 B)) st ill takes the advantage of the large vertical displacement at the tip of the frames, but it uses three bimorph beams to cancel the angular motion and two frames connected in between that rotate in the opposite directions to compensate the lateral shift on the platform. Moreover, the previous LVD actuator has two separate heaters for the two bimorph beams Thus a certain ratio of the driving voltages is needed to cancel the angular motion. The new design uses one uniform heater embedded along all actuators, so the driving is simplified and only one voltage is needed 1 2 ( l1) ( l2) ( L1) 3 ( l3) ( L2) Substrate Platform Bimorph I Frame II Bimorph III Frame I Bimorph II V + Substrate Platform Bimorph I Frame II Bimorph III Frame I Bimorph II V + 1 2 ( l1) ( l2) ( L1) 3 ( l3) ( L2) Substrate Platform Bimorph I Frame II Bimorph III Frame I Bimorph II V + Substrate Platform Bimorph I Frame II Bimorph III Frame I Bimorph II V + 1 2 ( l1) ( l2) ( L ) Frame Substrate Platform Bimorph I Bimorph II Lateral shift V2 + V1 + Frame Substrate Platform Bimorph I Bimorph II Lateral shift V2 + V1 + V1 + 1 2 ( l1) ( l2) ( L ) Frame Substrate Platform Bimorph I Bimorph II Lateral shift V2 + V1 + V1 + Frame Substrate Platform Bimorph I Bimorph II Lateral shift V2 + V1 + V1 + A B

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54 Figure 3 3 Schematic showing geometry design of the LSF LVD actuator The detail s of the LSF -LVD actuator design is shown in Fig ure 3 3 in which L1 and L2 denote the lengths of the two fram es, l1, l2 and l3 are the lengths of the three bimorph beams, and 1, 2 and 3 are the arc angles of the three bimorph beams denoting the sum s of the initial tilt angles and thermally actuated angles In order to keep the platform parallel to the substrate surface during the actuation, the arc angle s of the three bimorph beams should satisfy the following relation: 3 1 2 (3 1 ) According to Equation 2 10, the thermally -induced angle change is proportional to the bimorph len gth and temperature change Assuming a uniform temperature is obtained on all the bimorph beams the arc angles will be directly proportional to the bimorph lengths Thus Equation 3 1 becomes3 1 2l l l For simplicity, l1 and l3 are chosen as half of l2 to maintain a flat platform, i.e., 2 /2 3 1l l l (3 2 ) Th is also yields 2 /2 3 1 (3 3)

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55 The lateral shift of a level ed platform comes from five components: Actuator1, Frame1, Actuator2, Frame2 and Actuator3 as shown in Fig ure 3 3 Assuming each bimorph beam ha s a uniform curvature, the lateral shift of each component can be expressed as 1 1 1 1 1/ ) (sin l l LSactuator (3 4 ) 1 1 1 1cosL L LSframe (3 5 ) 2 1 2 2 2/ ) (sin 2l l LSactuator (3 6 ) 1 2 2 2cosL L LSframe (3 7 ) 3 3 3 3 3/ ) (sin l l LSactuator (3 8 ) By plugging Equation 3 2 and Equation 3 3 into E quation 3 4 ~ 3 8, and noting that the lateral shifts from the Actuator1, Actuator3 and Frame1 are opposite to the lateral shifts from the Actuator2 and Frame2, the total lateral shift of the platform can be expressed by : ) cos 1 )( (1 2 1 3 2 2 1 1 L L LS LS LS LS LS LSactuator frame actuator frame actuator (3 9 ) Thus the lateral shift and the tilting of the platform can be eliminated by simply choosing the same length for the two frames and the lengths of the bimorph beams to satisfy Equation 3 2 To maintain a uniform temperature over the three actuat ors, the embedded Pt heater is patterned all along the beams and the frames, so that a uniform heating can be obtained simultaneously by using a single driving voltage. To verify the LSF LVD design, thermomechanical simulations of the micromirror were con ducted using the MemMech -MemEtherm FEM co -solver in Coventorware [149] In order to simulate the initial reset position due to the residual stresses, the ambient temperature of the simulation was set to 150 K which gave the approximate residual stresses expected in the material layers. This does not affect the thermally induced actuation which is dependent on the

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56 temperature change It is not a problem to electrot hermalmechanical simulation either because the major heat dissipation mechanisms (convection and conduction) are not ambient temperature dependent [150] Fig ure 3 4 shows the simulation results of the vertical displacements at temperatures of 150 K (represents the initial reset position due to residual stresses ), 240 K and 340K respectively. The maxi mum lateral shift is only about 15 m for the entire vertical displacement range of 1116 m, with the maximum tilting angle of about 0.73. Figure 3 4. FEM simulation results of a LSF -LVD actuator s vertical displacements at different temperatur e by Coventorware A) 150K B) 240K and C) 340K 3 2 LSF -LVD based Tip -Tilt -Piston Micromirror 3. 2 1 Device D esign The LSF LVD electrothermal actuator design solves the lateral shift proble m in the two bimorph LVD design and also inspires new designs for transverse scanning devices as well. This section presents a new TTP micromirror design that is capable of not only large vertical displacement scan but also a cent er -stationary 2 D scan. Figure 3 5 shows the layout of the design where a squared mirror plate is connected by four sets of identical LSF LVD actuators at each side in a symmetric configuration. Each set of the actuator consists of two rigid frames and three bimorphs connected in series. Slightly different from the design shown in Figure 31, each actuator set is folded in a meandered shape. A B C

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57 By actuating the four identical actuators to maintain equal vertical displacements simultaneously a lateral -shift -free and tilt -free piston motion can be generated at each side of the mirror plate and the whole mirror plate as well T his symmetric structure also helps to further compensate the residual tilting caused by the temperature difference and fluctuation on the bimorph beam s. Tip tilt motions can be obtained by simply driving one or two actuators to pull down one side or two neighboring sides of the mirror plate. Therefore, by differentially driving the two pairs of opposite actuators, a center -stationary 2 D dynamic scan in the two orthogonal directions can be achieved. Figure 3 5 Top view layout of LSF LVD based TTP micromirror. A) micromirror with four LSF -LVD actuator and B) meandered LSF -LVD actuator. Both the piston and the tip tilt motio ns have been verified by electrothermomechanical simulation s using Coventorware. The mirror size of t he 3 D model is 0.5 mm by 0.5 mm the bimorph lengths of each LSF LVD actuator are respectively 100 200 and 100 and the two frames are 200 lo ng Figure s 3 6 A ) and B) show the mirror plate moved downward from its initial elevation at an ambient temperature of 150K with all four actuators being applied Mirror plate Bimorph beams Pt heater Frames B A

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58 by equal voltage s of 1 V Fig ure s 3 6 C) and D ) give the tip tilt motion results with a 1 V vo ltage applied to only one actuator and two adjacent actuators respectively. Figure 3 6 FEM simulation results of LSF -LVD TTP micromirror by Coventorware. A ) Displacement contour at 150 K to mimic the i nitial elevat ion B ) wit h 4 actuators excited with equal voltages C) 1 actuator excited and D ) 2 adjacent actuators excited. Table 3 1 FEM simulation results of TTP micromirror design by Coventorware Driving mode Displacement on mirror plate from initial elevated position m) Tilt angle of the mirror plate (degree) X Y Z Piston mode by 1 V on four actuators 0.384 ~ 0.384 0.395 ~ 0.395 128.39 ~ 128.40 1.146E 3 Tip Tilt mode by 1 V on one actuator 0.62 ~ 3.21 1.33 ~ 0.49 62.22 ~ 10.63 5.92 Tip Tilt mode by 1 V on two adjacent actuators 3.4 ~ 3.17 2.27 ~ 2.21 113.1 ~ 22.5 7 .3 6 A B C D

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59 The simulat ion results of the X, Y and Z displacements on the mirror plate are listed in Table 3 1, from which the tilt angle of the mirror can be estimated from the maximum di fference of the Z displacement over the entire mirror plate One can see that in the piston mode, the lateral shift in X and Y directions and the tilt angle of the mirror plate are further minimized by the symmetric four actuator structure compared to the single actuator results in Figure 3 4 3. 2 2 Device Fabrication A wafer -level combined surface / bulk -micromachining batch process has been developed for the device fabricat ion Al and SiO2 are selected for the bimorph materials and Pt is used for the hea ter material The process flow is outlined in Fi gure 3 7 it starts from patterning the Pt heater (0.2 by a sputtering and lift -off process on a 1 SiO2 coated wafer A ) An isolation PECVD SiO2 layer (~0.1 is then deposited B), fol lowed by an Al e -beam deposition (~1 and lift off for Al/SiO2 bimorph beams and the mirror surfac e C). A SiO2 plasma etching is performed to define the bimorph beams D ). Next, a backside SiO2 etch followed by a DRIE (deep reactive ion -etching) silicon etch is used to form a 20 40 crystal silicon) membrane for supporting the frames and the platform E ). After that, the wafer is diced and the separated dies are ready for front -side release. From the front side, a deep silicon trench e tch is performed to etch through the SCS membrane F ), and then the silicon underneath the bimorph beams is undercut by a Si isotropic etch, which releases the entire microstructure G ). Fig ure 3 8 shows SEM pictures of some fabricated devices. Two types of devices with different sizes have been designed and successfully fabricated. A 4 4 mirror array of type I device was fabricated as shown in Figure 3 8 C). The structure parameters of the devices are listed in Table 3 2.

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60 Figure 3 7 Cross -sectional view of f abrication process flow for a TTP micromirror Table 3 2 S tructure p arameters of fabricated TTP micromirrors Device Type Beam length Frame length Platfo r m Initial elevation I 100, 200, 100 200 500 187 II 150, 300, 150 500 800 646 B) : PECVD dielectric SiO 2 C) : Mirror SiO 2 etch D) : Al evaporation & l ift off F) : Backside SiO 2 & Si etch E) : Front side SiO 2 etch A) : Pt sputtering & l ift off G) : Front side Si anisotropic etch H) : Front side Si isotropic etch Si SiO 2 Pt Al

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61 Figure 3 8 SEM p ictures of fabricated TTP micromirrors. A ) Device Type I inset sho win g details of the actuator, B ) device Type II and C ) a 4 4 array of Type I device 3. 2 3 Experimental Results 3. 2 3 .1 Piston motion Before the characterization of a fresh device, an initial burn in phase is needed for the self annealing of the metal layers in the bimorphs, after which a good driving repeatability can be obtained The maximum voltage for safe operation is experimentally determined when an overheated burn pattern starts to develop on the bimorph Al layer at the actuators maximum temperature point, which is about at the center of the second bimorph beam The initial burn -in phase is then performed by applying an ac voltage of Vp( 0.5 + 0.5 200um Al/Pt/ SiO 2 500um SCS Meshed SiO2 500um 2mm A B C

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62 slightly less than the maximum voltage for at least a thousand driving cycles The initial curling angles of the bimorphs will increase after the burnin process resulting in a s light increase of the mirrors initial elevati on The static response of the devices piston motion ha s been characterized by using an Olympus BX51 microscope and a QC200 geometry measuring system. The mirror plate of the LSF -LVD TTP micromirror is displaced downward with a. DC voltage applied to the four actuators simultaneously By tracking the coordinates of the focal point on the platform through the microscope, the vertical displacement, lateral shift and the tilting of the mirror plate can be precisely measured Fig ure 3 9 A) shows the voltage -ve rsus vertical displacement plots of the two device types with two different sizes. Large vertical displacements of 214 and 621 have been obtained for a Type I and a Type II device at the driving voltages of only 4 V (171 mW ) and 5.3 V (184 mW ), respectively As can be seen there is almost no hysteresis between the downward and upward actuation. Good linearity (<2%) over a 400 range has been observed for Type II device. The testing result s from the devices of two different sizes also demonstrate that the actuation range can be easily scaled up by simply scaling the structure parameters of the actuator, so that larger vertical displacement in millimeter ranges are achievable based on this a ctuator design. The lateral shift and tilting of a Type II device are also measured and plotted in Fig ure 3 9 B) and C) The maximum lateral shift at the platform center is about 10 m which is only 1.6% of the 621 m platform size In addition, Figure 3 9 C) shows the maximum tilting angle is about 0.7 Considering the measured 0.2 tilting of the device package, the actual t ilting angle of the mirror plate to the device substrate over the entire actuation range is less than 0.5 The small lateral shift and tilting

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63 are believed to be mainly caused by mismatches among the four actuators due to fabrication process variations an d by the temperature fluctuations during the actuation Figure 3 9. Piston motion static response of TTP micromirrors. A) v ertical displacement versus applied voltages of both devices B) lateral shift and C) tilting of the mir ror plate during actuation for a Type II device Figure 3 10. Microsopic pictures of a Type II device at different applied voltages A) 0 V B) 3.5 V and C ) 5.5 V. B A C B A C

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64 Figure 3 10 shows the micros copic pictures of a T ype II device at applied voltages of 0 V A ), 3.5 V B ) and 5.5 V C) with the microscope focused on the mirror plate. The electrical resistances of the Pt heaters embedded in the a ctuators increase with the applied volta ges due to the t emperature coefficient of resistivity (TCR) of the thin -film Pt heater. Fig ure 3 1 1 A ) shows the electrical resistances of the Pt heaters (four in parallel) versus the vertical displacement (device Type II), in which a good linear correlation has been observed. This experimental relationship can be used for controlling the vertical position by monitoring the electrical resistances of the Pt heaters. Furthermore, by tracking the electrical resistances of the heater with a kno wn TCR of the Pt heater, the average temperature of the heater in the bimorph beams can be estimated. The TCR of the Pt heater has been measured to be 0.0029/C using a temperature controlled oven. The open -circuit, room temperature (23C) electrical resis tance of the four identical Pt heaters in par allel is measured to be 93 ure 3 1 1 B) shows the estimated temperature versus the applied electrical power, which offers important information for further characterizing the electrothermal performances of the actuator, such as the thermal resistances and the convection coefficient of the actuator. Figure 3 11. Experimental results of Type II device. A ) Vertical displacement versus heater resistance. B) Estimated temperature versus electrical power. B A

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65 3. 2 3 .2 Tip -tilt motion The tip -tilt motion has been verified using a Zygo interferometer Fig ure 3 12 A ) and B) show the interference patterns with one actuator and two adjacent actuators excited, respectively, both of which are consistent with the FEM simulatio n results in Fig ure 3 6 C) and D ). Figure 3 12. Zygo interferometric measurement A ) one actuator being excited, and B) two adjacent actuators being excited. An experimental setup with a laser beam incident on the mirror surf ace and voltage applied to four actuators individually has been used to measure the static scanning angles versus applied voltage on each actuator As shown in Fig ure 3 13, over 3 0 optical angles can be obtained from each actuator at 4.3 Vdc for a T ype I device and at 5.5 Vdc for a T ype II device Figure 3 13. Static tip tilt angles versus applied voltages on individual actuators of TTP micromirrors. A) T ype I device and B ) Type II device B A A B

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66 Dynamic 2 D scanning was performed by using the same experimental setup for the static scan By driving the two sets of opposite actuators with two pairs of ac voltages, which are respectively the two opposite sides of the mirror plate in each direction can be excited alternatively with a fixed 180 phase shift Thus, an op tical scanning range of more than 30 can be obtained independently in both axes with a stationary mirror center Fig ure 3 14 A ) and B) show the scanning patterns with 0 and 90 phase difference of the two ac voltage pairs for both directions By sweepin g the frequency, the measured resonant frequency of the scanning mode was about 2.3 kHz Fig ure 3 14 C) shows a 1530 raster scanning obtained by driving one axis at resonant frequency (fast scan) with Vp = 1.5V and the other at 47 Hz (slow scan) with Vp = 3.5V Scanning patterns by two adjacent mirrors from a 4 4 mirror array is also demonstrated in Fig ure 3 14 D ). Figure 3 14. 2 D scanning patterns of TTP micromirrors A ) Line scan by 0 phase difference of the two axes B ) Circular scan by 90 phase difference. C ) A raster scan pattern. D ) Scanning patterns by two adjacent mirrors from a 4 4 mirror array. 15 o 15 o 30 o 15 o Bonded mirror array Scanni ng pattern from two adjacent mirror B A D C

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67 3. 2 3 .3 Dynamic r esponse The frequency responses of both types of devices have been measured using a Polytec OFV511 las er Doppler vibrometer. W ith an applied voltage of (1+ 0 .5sin t) V, a s shown in Fig ure 3 15 A ) and B) t he first resonance peaks are observed at 1.334 k Hz and 453 Hz for T ype I and T ype II devices respectively Fig ure 3 15 C) and D ) show the measured displacement contour on the mirror plate at the first two resonance peaks of a T ype II device, demonstrat ing that the first resonance mode is a vertical mode and the second corresponds to a rotational mode. Figure 3 15. Frequency response measurement of TTP micromirrors by a laser vibrometer. A ) Device Type I B) device T ype II C) piston mo tion at the first resonance peak and D ) rotational motion a t the second resonance peak. The devices demonstrate a mechanical resonance shift at different DC bias, this is believed to be related to the variations of the bimorph materials mechanical properties at different temperature The temperature on the bimorp h beams increase with the applied DC bias, and the mechanical property, i.e., the Young s moduli of the bimorph materials, especially the Al, B D C A

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68 decrease accordingly [151, 152] therefore, resulting in a degradation of the actuator s stiffness. The piston mode resonance frequenc y shift of a modified T ype II device was characterized using the same laser Doppler vibrometer. The device has a mirror size of 1 mm and thicker Si layer of 40 m underneath the frames and the mirror plate. Figure 3 16 shows the measured r esults at voltages of ( Vdc + 0 1 sin t) V with the DC bias Vdc changing from 1 V to 5 V. A maximum resonance shift of 26.9% was observed with the DC bias being increased from 1 V (398 Hz) to 5 V (291 Hz). Table 33 lists all the measured resonance frequencie s of the piston mode at different DC bias, the ir neutral mirror elevation s and the estimated average temperature on the bimorphs. Figure 3 16. Measured piston resonance s of a modified T ype II device at differen t DC bias Table 3 3 Measured piston resonance, DC vertical displacements and estimated average temperature on the bimorph at different DC bias, for a modified T ype II device DC bias (V) Resonance frequency (Hz) M irror elevation ( m) Estimated average temperature (C) (ambient temperature 27 C ) 1 398 591.4 49.2 2 375 451.6 101.3 3 346 282.8 157.1 4 316 155.3 203.6 5 291 35.5 231.3

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69 3 3 LSF -LVD Microlens Scanner 3. 3 .1 Device Design and Fabrication As discussed in section 1.2.3 miniature microlens scanners capable of millimeter range large vertical scan are highly desired for the focusing mechanism in biomedical imaging endoscopes. The same LSF LVD electrothermal actuator design presented in the previous section can be di rectly applied to microlens scanner s As shown in Figure 3 -1 7 instead of an Al coated mirror plate in the TTP micromirror design, a rigid hollow frame can be fabricated to hold a microlens on top By driving the four actuators simultaneously the piston motion of the lens holder can be excited T hus the focal plane of the microlens can be vertically displaced with minimized lateral shift and tilt on the microlens s optical axis. Figure 3 17. Schematics of LSF LVD lens scanner d esign. A ) Top view layout and B) cross sectional view. The lens holder shares the same fabrication process as that for the micromirror except that a n opening is formed on the silicon platfor m instead of Al coating for the mirror surface The microlens ca n be supported by a transparent meshed SiO2 structure patterned above the silicon platform or through the buried oxide layer by using an SOI wafer. In the first case, the Si underneath the meshed SiO2 can be removed during the final release step (Figure 3 7 H )), so a hollow platform with transparent SiO2 for supporting a mico lens can be formed In the second A Micro lens Bimorph beams B Microlens Frame Substrate

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70 case, an SOI wafer with a thin device layer (about 40 m) is employed. With the buried SiO2 holder a transparent reservoir is formed for holding the polymer lens. Devices by both methods have been fabricated with the SEM photos shown in Figure 3 1 8 Figure 3 18. S EMs of fabricated LSF LVD lens holder A ) w ith meshed SiO2 on top of Si platform B) w ith buried SiO2 membrane from SOI wafer and C) a lens hold er with Polymer lens integrated As a demonstration of the proof of concept, a polymer microlens is formed on a lens holder by the liquid dispensing method as previously used by Jain et al. [153] P olymer droplets (photoresist NR7 100) w ere precisely dispensed onto the transparent reservoir o f the lens holder using a nanoliter injection system and then baked at a 120 C for 30 minutes to form the B A C 1 mm Polymer microlens Meshed SiO2 1 mm Buried SiO2 membrane 1 mm

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71 micro len s by the surface tension. An SEM picture of a polymer microlens formed on the lens holder is shown in Figure 3 1 8 C). 3. 3 .2 Experimental Results An imaging experiment was performed to test the image quality of the integrated polymer lens. Fig ure 3 1 9 A ) s how s the schematic of the test setup T he image generated by the diameter polymer micro lens is shown in Fig ure 3 1 9 C), where the resolution of about 5 be resolved without distortion. The focal length is measured to be about 2 mm with a depth of focus about 0.2 mm. Fig ure 3 1 9 Lens scanner i maging experiment A ) Schematic of the imaging experiment B) p hotograph of the mask and C) i mage of the test pattern generated by polymer micro lens T he LSF LVD actuation w as characterized by applying a single DC voltage to all the four actuators simultaneously. Large vertical displacements of 730 m have been obtained at only 7.5 V dc (323 mW) f rom a same device before and after the polymer lens being integrated respectively, as shown in Fig ure 3 20 A ). Figure 3 20 B) show t he measured lateral shift and tilt angle. The maximum lateral shi ft at the center of the polymer lens is only 13 (<2% of the vertical scan range) and the maximum tilt angle is 0.74 (also the initial tilt angle) A B C UFUF = 633nm fobjfCCD20X UFUF = 633nm fobjfCCD20X 10 m 3 0 0 m

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72 during the entire actuation range. The resonance frequency of the purely -vertical motion mode is measured to be 761 Hz and 488 Hz before and after the lens is integrated. Fig ure 3 2 1 shows a series of microscopic pictures (top view) of an LSF LVD microlens at different actuation voltages, where no obvious lateral shift was observed. Fig ure 3 20. Lens scanner e xperimental results A ) Vertical displacement versus applied voltages and B) l ateral shift and tilting of the lens during the actuation. Fig ure 3 2 1 Mic ro scopic pictures of the microlens at diff erent applied voltages. A ) 0 V, B) 4 V and C) 6 V. 3 4 Lumped Element Models and Improved LSF -LVD Actuator Design As demonstrated in previous sections, large vertical displacement with small lateral shift and tilting is achieved by t he LSF LVD actuator an d versatile optical scanners are developed based on this novel design. However, the original design suffers from a slow thermal response resulting in a significant decay of the achievable scan range at low frequency below 100 Hz as A B C A B

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73 can be seen in Figure 3 15. To further study this behavior, an electrothermal analysis of the actuators steady -state and transient response is performed in this section An improved LSF LVD actuator design with Si -free frames for faster thermal response will also be presented. 3.4.1 S teady S tate Analysis To analyze the steady -state response, the actuators initial elevation, as the reference from where the downward actuation starts, needs to be considered. The bimorph beam is initially curled up after device release due to the residual stress of the thin films, resulting in an initially elevated and flat position of the connected platform This initial elevation is determined by the initial arc angles of the bimorph beams and the length of the frames, as can be expressed by: 0 0sin 2 L H (3 10) Assuming an average temperature increase T is introduced to all the three bimorphs by applying a voltage V to the embedded heater [137] : 2 / ] 1 4 1 [0 2 E TR R V T (3 11) where 0 ER and are respectively the heater electrical resistance at initial temperature and its temper a ture coefficient and TR is the equivalent thermal resistance of the actuator structure As shown in Figure 3 2 2 a net downward piston displacement of the connected platform that is generated by the bimorph actuation with a thermally induced arc angle of the first bimorph T can be expressed as: ) sin cot cos 1 ( )] sin( [sin 20 0 0 0 T T TH L H (3 12) Therefore, by combining Equation 3 1 1 Equation 2 10 and Equation 3 12, an expression of th e piston displacement of an LSF -LVD actuator directly related to the applied voltage on the heater can be obtained as following:

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74 ]} 2 / ) 1 4 1 ( sin[ cot ] 2 / ) 1 4 1 ( cos[ 1 {0 2 0 0 2 0 E T b b E T b bR R V t l R R V t l H H (3 13) Fig ure 3 2 2 Schematics showing actuation principle. A) a single bi morph and B) an LSF LVD actuator. T he actuators equivalent thermal resistance RT may have large variations due to the convection, packaging thermal contact and fabrication variations, so it is usually hard to be precisely estimated. Fortunately, RT is th e only parameter with the most uncertainty and can be characterized from the experimental results Figure 3 2 3 shows the experimental and the calculated results from Equation 3 12 of a Type II device with different values of RT. By fitting RT to the experimental results, it was found that t he equivalent thermal resistance to be about 1.08104 K/W which yields a minimum average difference of 3.7 % between the experimental and calculated results in 0 3 V actuation range T he higher order effe cts such as higher order temperature coefficients of the heaters electrical resistance, heat dissipation through radiation and the mechanical properties variation s of the bimorph materials at high temperature are not considered in the model. The nonunif ormity of the temperature over the bimorphs is also neglected. The combination of all these neglected effects causes the more significant deviation Temperature increase A Plat form Bimorph I Bimorph II Bimorph III Frame I Frame II B H H 0

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75 between the calculated results and experimental results at high actuation voltage, i.e., the high temperatur e range, as can be seen in Figure 3 2 3 However, this simplified model can be used to estimate the uncertain properties, and well predicts the device behavior in about 60% of the entire actuation range. Figure 3 2 3 Applied vo ltage versus vertical displacement for a Type II device with experimental and calculated results by different values of equivalent thermal resistance. 3.4.2 Transient Thermal Analysis Figure 3 2 4 shows the equivalent circuit of a simplifie d transient lumped element model. Figure 3 2 4 The equivalent circuit of a simplified transient lumped element model of LSF -LVD actuator sC mC hT camRcasRamTam R sR mR asCamCasT0Tas R Q

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76 Table 3 4 Parameter s used in the lumped element model LSF LVD actuator Parameters Description T h Tas Tam Ras Ram Rs Rm Rcas Rcam Cas Cam Cs C m The maximum temperature on the actuator at applied current I The temperature at the substrate end of the actuator The temperature at the mirror end of the actu ator Thermal resistance by conduction of the actuator from Th to Tas Thermal resistance by conduction of the actuator from Th to Tam Thermal resistance from the substrate end of the actuator to T0 Thermal resistance from the mirror end of the actuator to T0 Thermal resistance by convection on the actuator from Th to Tas Thermal resistance by convection o n the actuator from Th to Tam Thermal capacitance from Th to the substrate end of the actuator Thermal capacitance from Th to the substrate end of the actuator Thermal capacitance of the device substrate Thermal capacitance of the mirror plate The three regions, i.e., the actuator, device substrate and mirror plate are lumped by their equivalent thermal resistances and capacitors The lumped par ameters are listed in Table 3 4 The actuator is split into two parts from its maximum temp erature hT location [137] and the heater is lumped at the maximum temperature point Therefore, the heat transfer equation of the actuator region can be expressed as: Q R T T R T T R T T R T T dt T T d C dt T T d Ccam h cas h am am h as as h am h am as h as ) ( ) ( ) ( ) ( ) ( ) (0 0 (3 14) w here Q is the heat generation i.e., the applied power by Joule heating To simply derive the time constant of the maximum temperature increase hT ( 0T Th ), s everal approximations are made First, the time derivatives of asT and amT are much smaller than that of amT : dt dT dt dTh as dt dT dt dTh am (3 15) T hus they are neglected considering the large thermal capacitances of the device substrate and the mirror plate associated with their th ermal time constant s Second, the relationships of asT amT and hT from the steady -state solutions are used in the heat conduction terms:

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77 ) ( ) (0 0T T R R R T Th as s s as (3 1 6 ) ) ( ) (0 0T T R R R T Th am m m am (3 1 7 ) Suppose a sinusoidal power input is applied as the heat generation. Then Equation 3 1 4 reaches: t j T h h ae P R T dt T d C (3 1 8 ) where aC is the actuators thermal capacitance, TR denot es the equivalent thermal resistance which can be experimentally characterized as shown in section 3.4.1 and T aR C is defined as the time constant A general solution of Equation 3 1 8 can be obtained: ) ( / 1 / ) (/ t t j a he e jC P t T (3 1 9 ) The 3dB cut -off frequency of the longtime ( steady -state) solution can be found as: 2 1off cutf (3 20) Figure 3 2 5 Measured frequency respo nse of a Type II device 3 dB Thermal cut off ~10 Hz

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78 The calculated results of a Type II device ( with 1 1 mm2 mirror plate) show the time constant of 17.3 ms and the corresponding 3dB cut off frequency of about 9.4 Hz which is reasonable compared to the measured cut off frequency of 10 Hz, as shown in Figure 3 2 5 3.4.3 Improved LSF -LVD Actuator Design with Faster Thermal Response As discussed in the previous section the transient response of the LSF LVD actuator is determined by the time constan t, i.e., the thermal capacitance and the equivalent thermal resistance of the actuator structure Therefore, in order to improve the scanning speed, i.e., the cut off frequency, the actuator structure needs to be modified with reduced thermal capacitance a nd resistance. 3. 4 3 .1 Reducing the T hermal R esistance The original LSFLVD actuator design, as presented in section 3. 2, uses pure SiO2 at the substrate anchoring point This is to tak e the advantage of the materials low thermal conductivity for the pur pose of a good thermal isolation and heating efficiency Therefore, as one solution to improve the transient response, one can reduce the equivalent thermal resistance by extending the top Al layer of the bimorph to the device substrate The good thermal c onductivity of the Al layer makes it an additional thermal path between the actuator to the bimorph, thus can efficiently reduce the global thermal resistance On the other hand, the brittle SiO2 layer is easy to fracture and make the device very fragile t o external shock. The good elasticity of the Al metal layer helps improve the device reliability and sustainability to certain external shock s T he trade off here is the higher power consumption. Figure 3 2 6 shows the SEM picture of a Type II device with an additional Al metal layer on top of the pure SiO2 thermal isolation region in the original design (Type III). The equivalent thermal resistance of the new device was experimentally characterized by the same way as in

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79 se ction 3.4.1 and the measured value was 4.5 103 K / W With the thermal capacitance having 40 the transient model predicts a reduced time constant of 13 ms resulting in a cut off frequency of 1 2 Hz, which is in good agreement with the measured result of about 13 Hz as shown in Figure 3 2 7 Figure 3 2 6 SEM picture of the modified LSF -LVD actuator with additional Al layer on the substrate junction. Figure 3 2 7 Measured frequency response of a Type III device 3 dB Thermal cut off ~ 13 Hz 395 Hz Additional Al 100 m

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80 Due to the smaller thermal resistanc e, the consumed power for a Type III device to reach 600m vertical actuation is 389 mW, about 2.5 times of the power consumed by a Type II device to reach the same actuation, i.e., the same temperature change on the actuator This result is in a good agreement with the value of 2.4 which is ra tio of the characterized equivalent thermal resistances of the two devices, considering that the thermal resistance and the actuation power are in scale for the same temperature actuation. A drop test is performed to assess the improvement on sustainabilit y to external shock of Type III device Both Type II and Type III devices are packaged into a dual in line package (DI P) that is plugged in a small breadboard with a total mass of about 5 0 g The whole pac kage wa s dropped to concrete floor starting from 2 inches height with an increase step of 2 inches The Type II device with the pure SiO2 substrate anchoring completely brok e on all the four substrate anchoring junctions and the mirror pla te fell off the device substrate at the height of 6 inches The Type III device was found to be much more sustainable to external shock. The first slight fracture on the edge of one actuator s substrate anchoring junction developed at the height of 14 inches with the four actuators still operational, and th at actuator was completely broken at the height of 38 inches 3. 4 3 2 Reducing the T hermal C apacitance The thermal capacitance comes from all the materials composing the actuator structu re The majority of the thermal capacitance in the original design is the thick Si layer underneath the two frames as shown in Figure 3 8 The two frames are used to amplify the vertical actuation and compensate the lateral shift The employment of bulk S i layer is to maintain t he frame flatness during thermal actuation This layer of Si usual ly has a thickness of 2545 m and is formed by a DRIE step on a regular Si wafer as described in section 3.2. A n SOI wafer can also be employed to take the advantag e of the accurate and uniform thickness of its device layer

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81 T he thermal capacitance can be significantly decreased if removing this Si layer, to be about one sixth of the original design with 25 m thick Si frames, resulting in a considerable improvemen t on the thermal cut off frequency that can be increased by six time s This sub section presents a new LSF LVD actuator design with Si -free multilayer thin film frames The multilayer thin film (or multimorph) structure has optimized thickness ratios to c ompensate the thermally induced strain so that the flatness can be maintained during the thermal actuation In this design, the points of interest are the flatness, i.e., the radius curvature of the multimorph, and the conditions for a flat multimorph st ructure with its flatness insensitive to temperature variation, i.e., the thermal stress In the original LSF -LVD actuator structure, the three bimorph segments are actually composed of multiplayer thin films, including the two major bimorph materials, the bottom SiO2 and top Al layer, as well as a thin layer of Pt as the heater material and thin SiO2 and Cr metal layers as the dielectric and adhesion layer s The thickness of the heater, dielectric and adhesion layers are much smaller than the major bimorph actuator layers and thus can be neglected in the analysis of bimorph actuation. To obtain a flat and temperatureinsensitive multimorph structure that replaces the bulk Si -supported frames in the original design, an additional layer of a material that has TCE smaller than Al needs to be added on top of the current bimorph structure SiO2 is a good candidate material of that additional layer for the convenience of the current fabrication process Therefore, the problem is reduced to simply derive the thickn ess of this additional top SiO2 layer that is needed to compensate the thermally induced strain and maintain the flatness of the multimorph structure, taking that thickness of all other layers are kept the same as needed for the bimorph actuator. Figure 3 2 8 shows a side -view schematic of the multimorph structure.

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82 Figure 3 2 8 The side -view schematic of the multimorph st ructure An analytical model for the induced -strain multimorphs is developed by Garcia et al. in [154] This generic model predicts the deformed shape, the corresponding tip bending moment, as well as the position of the neutral axis (NA) and the composite bending rigidity for any types of induced strain s The multimorph s curvature radius derived from this model, as expressed in Equation 3 2 1 is used to analyze the new Si -free LSF -LVD actuator design. n i i N i n i i N i i iz z D z z C I E R1 1)] ( [ )] ( [ (3 2 1 ) iz is the position of the symmetry axis of each layer that can be calculated according to the thickness it as shown in Figure 3 2 8 : n i j j i it t z12 (3 2 2 ) Nz is the position of the neutral axis: n i i i n i i i i NA E A E z z1 1 (3 2 3 ) C and D are rela ted to the biaxial Youngs modulus, thickness and the stress of each layer: Dielectric SiO 2 Top SiO 2 (PECVD) Bottom SiO 2 (Thermal/PECVD) Cr Pt Cr Cr Al Z N Z 1

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83 n i i i n i i k i k i i i k i k i i iA E A E t t t A E t t t A E C2 1 1 2 1 2 1 1 2 1)] 2 ( [ 2 2 (3 2 4 ) ) ( ] ) [(2 1 1 2 1 1 n i i i n i i i i i i i iA E A E A E A E D (3 2 5 ) If the thermal stress dominate such as in electrothermal bimorph actuators: Ti i (3 2 6 ) T o reach a flat multimorph structure with its flatness insensitive to temperature variations as shown in Figure 3 28, the structure parameters of each layer are required to lead to an infinite curvature radius Therefore, the required thickness of the top SiO2 can be easily calculated by solving Equation 3 2 7 given that the thickness es of all the other layers are fixed. 0 )] ( [ )] ( [ )(1 1 1 1 1 n ii N i n i i N i i iz z D zz C I E t R (3 2 7 ) Figure 3 2 9 T he calculated reciprocal of curvature radius versus the top SiO2 (PECVD) thickness

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84 Figure 3 2 9 shows the calculated results of the reciprocal of the curvature radius versus the thickness of the top SiO2 (PECVD), for a thermal bottom SiO2 structure at the thermal ly induced stress by a temperature change of 300 K The material and structure para meters used in the calculation are listed in Table 3 5 As can be seen in Figure 3 2 9 there is one solution for t1 at 1.509 m be sides at infinitely large that gives a n infinite radius curvature. The bottom SiO2 can be also fabricated by PECVD, which le ads to a slight change of the required top SiO2 thickness Table 3 5 Material and structure p arameter s used for the analytical calculation [155] Layer # Material T hick ness ( m) TCE ( 107K1) Young s Modulus (GPa) Poisson Ratio Specific Heat 103JKg1K1 Densit y ( 103Kg m3) 1 PECVD SiO 2 24.125 73 0.175 0.84 1.961 2 Al 1 248.333 68.85 0.36 0.9 2.692 3 Cr 0.02 29.0554 160 0.42 0.447 7.19 4 PECVD SiO 2 0.2 24.125 73 0.17 5 0.84 1.961 5 Cr 0.01 29.0554 160 0.42 0.447 7.19 6 Pt 0.2 90 146.9 0.35 0.133 21.5 7 Cr 0.02 29.0554 160 0.42 0.447 7.19 8 Thermal SiO 2 1 5 67 0.15 0.84 2.353 FEM analysis by Intellisuite [155] has also been performed to verify the analytical calculation. A 500long cantilevered multimorph structure, as in Figure 3 30 A ), is built with the same structural parameters as in the above analytical calculation and a variable thickness for the top SiO2 layer The default temperature for an ideally flat structure is 0 C A thermomechanical analysis is then performed with the temperature changed to 150 and 300 C. T he simulation result s of the cantilevers tip deflection s after the temperature change can be used to assess the curvature radius, i.e., the flatness of the structure with different top SiO2 thickness es The optimized top SiO2 thickness can be found once a zero tip deflection is reached. Figure 3 30 C) shows the tip deflections from the simulation results for both thermal bottom SiO2 and PECVD bottom SiO2

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85 structure, and the optimized top SiO2 which are consistent with the analytical calculations. Figure 3 30. FEM simulation results of the multimorph structure. A ) Deflection co ntour of the SiO2 (thermal bottom SiO2) at 300 C, B) Tip deflections at different SiO2 thickness, C) Si -free LSF LVD actuator, at temperature of 150 C to mimic the initial elevation. Devices based on the new Si -free LSF LVD actuator design are fabricated and characterized. An additional PECVD SiO2 deposition is added in the fabrication process described in section 3.2 after the Al layer to form the top SiO2. A meshed pattern is formed on the two frame segments so that the Si underneath the frames can be completely etched together with the Si underneath the bimorph beams in the final release step. Figure 3 -3 1 shows the SEM B A C

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86 pictures of the fabricated device (Type IV ) with 1 bottom SiO2 and 1.52 top SiO2. Because the residual stress of the thin films mainly comes from the thermal stress, the initially released multimorph frames also demonstrate a good flatness. Figure 3 3 1 SEM pictures of a TTP micro mirror (device Type IV ) by the Si -free LSF LVD actuator A) the entire device, B) Si -free LSF -LVD actuator and C) a close -up of the bimorph and multimorph junction. As can be seen from Figure 3 2 9 the flatness of the multimorph frame is very sensitive to the thicknesses of the top SiO2, and the fabrication variations may easily affect the frame flatness. But fortunately, thanks to the symmetry of the actuator structure, the tilting and lateral shift caused by the curling of the two frames are in opposite directions and still cancel each other. Therefore, the lateral -shift and tilt -free vertical actuation can still be guaranteed even the two frames are not ideally flat. Based on the simulation results in Figure 3 30, a 4.6 % variation of the targeted top SiO2 thickness induces only about 1.7 % decrease on the vertical displacement for the actuation by a 200 K temperature change. An additional Al layer on the substrate anchoring junction as in Type III device is also applied to the Type IV device for reducing the thermal resistance, and it shares the same B 1 mm Multimorph frames C Bimorph Multimorph A

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87 structural parameters for the bimorph segments as the previous designs and thus should have the same DC actuation r ange for both piston and tip -tilt mode. The equivalent thermal resistance is experimentally characterized to be 5. 1 103 K / W by the same way as described in sub -section 3.4.2, resulting in a calculated thermal time constant of about 2. 2 ms and a significant ly improved cut -off frequency of about 71 Hz due to the smal l thermal capacitance of the Si free actuator. Figure 3 3 2 Measured frequency response of a Type IV device Table 3 6 Summary of the transient perf ormance of the device Type II III and IV Device Type T hermal resistance (W/K) Frame Si th ickness Calculated time constant (ms) Thermal cut off frequency (Hz) Power at 600 m piston actuation (mW) Resonance frequency (tip tilt driving ) (Hz) Calculated Measured II 10.810 3 20 17 9. 4 10 155 372 I II 4.5 10 3 40 13 12 13 389 395 IV 5. 1 10 3 0 2.2 71 78 311 355 As shown in Figure 3 3 2 the measured 3dB cut -off frequency is about 78 Hz which is reasonably close to the calculated result. Note that the resonant frequency is also decreased to 3 dB Thermal cut o ff ~ 78 Hz 355 Hz

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88 355 Hz due to the weakened stiffness of the Si -free LSF LVD actuator Table 3 6 summarizes the transient performance of the original design (Type II ) and the improved designs (Type III and IV ). 3 .5 Summary This chapter present s a novel LSF -LVD electrothermal actuator design and various optical micro scanners developed from this actuator design along with preliminary experimental verification. The fabricated TTP micromirrors have demonstrated over half millimeter vertical scan range on a 2 -mm device footprint with negligible lateral shift of less than 10 angle of about 0.5 The same micromirrors have also demonstrated a center -stationary 2 D scan with over 30 optical angles in both axes. Both piston and rotation motion require driving voltages lower than 6 V. An LSF LVD lens scanner with an integrated 600lens has also been demonstrated and the lenss focal plane can be vertically displaced up to about The devices steady -state and transient responses are analyzed using lumped eleme nt models, and an improved Si -free LSF -LVD actuator is designed employing both analytical and FEM models. The improved design demonstrates a fast thermal response with the thermal cut -off frequency up to 78 Hz. The se LSF LVD design s solves the lateral shif t problem, simplifies the driving for the piston motion, resolves the non -stationary mirror center scan in the previous LVD designs and improves the devices fill factor as well. The actuator design also shows great potential to achieve much larger vertical displacement and larger actuation force by simply scaling up the dimensions of the actuator.

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89 CHAPTER 4 FULL -CICURMFERENTIAL SCAN NING MICROMIRRORS AND L IGHT WEIGHT MICROMIRRORS As discussed in Chapter 2, for the endoscopic imaging in intravascular and i nternal organ applications such as in heart arteries and lung bronchi, a miniature full -circumferential scanning (FCS) mechanism is highly desired. In this chapter, a new FCS endoscope design by using a dual -reflective 1 D electrothermal micromirror with l arge rotation al scanning range is proposed. A novel process developed to fabricate the dual reflective micromirror will also be presented in this chapter. Furthermore, to guarantee the mirror flatness, the currently developed micromirrors use a layer of s ingle -crystal silicon (SCS) for the mirror plate. However, the mass of the SCS layer increases quadratically with the mirror aperture size. Thus increasing aperture size will reduce the scanning agility of micromirrors In this chapter, a self aligned deep trench process that can produce light -weight micromirror s with SCS ribs is proposed and some preliminary results are presented. 4 1 Full -Circumferential Scanning Endoscope Design Various FCS imaging probes have been reported [41, 47, 48] Early FCS research efforts focused on spinning the entire optical fiber and its assembly [41] which has the limitations of scanning speed and optical coupling stability Other reported FCS endoscopes employ commercially available mi cromotors to spin mirror s or wedge prisms to obtain 360 scanning [47, 48, 156, 157] MEMS micromirrors by electroastatic or piezoelectric actuations are not prac tical for the circumferential scanning because their achievable deflection angles are relatively small. An MEMS scratch drive array to generate a 360 rotation was proposed by Ayers et al. [158] ; however, due to its rather complicated structure and fabrication process, n o

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90 B further experimental results have been reported yet. On the other hand, electrothermal micromirrors have the advantages of much larger scanning angles than other actuation methods, and thus ha ve potential to realize full -circumferential scanning. Fig ure 4 1 The conceptual schematic of the FCS -EOCT imaging probe with a dual reflective MEMS mirror. A ) Two -fiber endoscope. B) One -fiber endoscope. Figure 4 1 A) shows the schematic of the proposed FCS endoscope. The full circumf erential scanning mechanism utilizes two light beams and a dual reflective 1 D rotational micromirror. Two optical fibers are used to deliver light beams to both surfaces of a 1 D rotational micromirror. The micromirror can rotate 45 (or 90 ) and thus a 180 optical scanning is obtained from each mirror surface, resulting in a full 360 scan. To avoid using two fibers which may require two imaging system s another 1 D MEMS mirror can be used as an optical switch as shown in Fig ure 4 1 B). The light beam i s directed to only one side of the dual reflective mirror when the first MEMS mirror switches between its on and off state, so the two half -circumferential ranges can be scanned and imaged alterna tively using one imaging system To realize the full -cir cumferential scanning from both incident beams, the scanning Fiber 1 Fiber 2 GRIN GRIN MEMS mirror 45 rotation A B

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91 micromirror must be reflective on both surfaces and capable of at least 90 (45) mechanical rotation angles. The design details of such a micromirror are discussed in next section. 4 2 Dual -Ref lective Micromirror 4.2.1 Device Design Electrothermal micromirrors using Al/SiO2 as bimorph materials with polysilicon heaters embedded have been developed and typical scanning ranges of about 40 can be obtained [52, 64] However, as mentioned in section 2.2.2, the embedded polysilicon heater exhi bits problems of hysteresis and self annealing effect s which limit their stable scanning ranges. Fig ure 4 2 The schematic of the dual reflective electrothermal micro mirror design A) Cross sectional view, and B) top view. To reach the larger deflection angle required for FCS scanning, a 1 D electrothermal micromirror using a Pt heater at the end of the bimorph beams instead of embedded along is proposed. The micromirror has reflective Al coating on both sides E ach side cover s a half of the full -circumferential scanning range. As shown in Figure 4 2 A ), the mirror plate is composed of a layer of single -crystal silicon for the mirror flatness, and is coated with Al on top of SiO2 on both sides of the mirror plate The mirror pla te is connected to the silicon substrate by an array of A Pt Heater Thermal isolation Thermal isolation Bimorp h beams Buried oxide Front side Mirror B Single crystal silicon Back side mirror

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92 Al/ SiO2 bimorph beams. A Pt layer is integrated at the substrate end of the bimorph beams for electrothermal actuation, covered by Al for good thermal conduction to the bimorph beams. The thin SiO2 la yer between Al and Pt is used for electrical isolation The bimorph beams are thermally isolated by meshed SiO2 regions at both ends for a high heating efficiency. The top view of the mirror design is shown in Figure 42 B). The bimorph beams curl up after being released due to the residual stress es of the thin films as shown in Fig ure 4 2 A) From the thermo -mechanical analysis discussed in section 2.2.3, the achievable actuation angle for a given bimorph length and thickness is proportional to the average temperature rise as shown in Equation 2 14. FEM electrothermomechanical simulation with CoventorWare has been performed to verify the theoretical calculation. The meshed model is a section of the complete micromirror structure with two bimorph beams conne cting the substrate and the mirror plate. The SiO2 thermal isolation at both ends. Convection was considered in the simulation with the convection coefficient of 170 W/ K m2 on the bimorph beams and 126 W/ K m2 on the mirror plate [150] Fig ure 4 3 FEM simulation results of a hingedbimorph 1 D mirror by CoventorWare A ) T emperature distribution at applied voltages of 60 mV and B) displacement contour at applied voltage of 8 0 mV. (Note: ambient temperature: 150 K) A B 150K Curled downward with about 88 rotation angle 350K 300K 200K

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93 Fig ure 4 4 T heoretical calculation and FEM simulation results of r otational angle versus average temperature change for a 1 D hinged-bimorph micro mirror design Figure 4 3 shows the temperature distribution and bimorph deflection at an applied voltag e of 60 mV. The average temperature on the bimorph region is about 91% of the maximum temperature at one end where the Pt heater is located as shown in Figure 43 A ). The average temperature versus the deflection angle from the simulation results was compa red with the theoretical calculation from Equation 2 14 in Figure 4 4, a maximum mismatch about 6% is observed. 4.2.2 Device Fabrication A combined surface / bulk micromachining process is presented in the previous chapter for fabricating single -side micromi rror s. In order to realize the reflective surface s on both sides of the mirror plate, a new process is developed for patterning a reflective coating on the backside of the mirror by using a self aligned deep trench silicon etching process.

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94 Figure 4 5 show s the cross -sectional view of the pr ocess flow. It starts from an SOI wafer with 1 thick thermal oxide on both sides. Fig ure 4 5 Cross -sectional view of the fabrication process flow for the dual reflective micromirror First Pt sputtering and lift-off is performed on the device layer side of the SOI wafer to form the heater pattern A ). Then a thin PECVD SiO2 layer is deposited, followed by Al e beam evaporation and lift -off for the bimorph and front -side mirror surface B) N ext, a SiO2 etch is performed to define the thermal isolation, bimorph beams and mirror patterns C). Next, a Si A ). Pt deposition and liftoff B ). Mirror SiO 2 etch and Al deposition /lift off H ). Buried SiO 2 etch I ). Front side release SiO 2 Pt Al F ). Back side Al deposition C ) SiO 2 etch D ). Backside SiO 2 etch E ). Backside Si DRIE G ). Back side Si isotropic etch Mirror plate Thermal isolation Bimorph beams Platinum heater

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95 backside SiO2 etch D ) is performed after the backside alignment and photolithography to form the opening for deep reactive ion etch (DRIE ) of backs ide Si. The backside Si DRIE stops at the buried SiO2 layer of the SOI wafer and defines the backside mirror region After that, another Al laye r is deposited on the buried SiO2 layer for reflective coating on the backside mirror surface F ), followed by a Si isotropic etch G ) and buried SiO2 etch H ) to separate the mirror plate from the substrate. Finally the device is flipped over for front -side Si DRIE, with first an anisotropic etch to separate the mirror plate from the substrate followed by an isotropic undercut of the Si underneath the bimorph beams, the mirror plate is released with a initial tilt angle due to the residual stress of the bimorph beams I). A few SEM pictures of fabricated device s are shown in Fig ure 4 6 a single -side d 1 -D micromirror with bimorph beam length 300 m and in A ) and a dual reflective micromirror with 200 m in B) The bimorph beams are 10-m wide with 8 m gap distance and the mirror plate is 1 mm by 1 mm. The mirror surface quality was measured by a Wyko NT100 white ligh t optical profilometer Figure 4 7 shows line scans on both mirror surface, t he measured radii of curvature for both the front and back mirror surfaces are about 129 mm and 1 32 mm, respectively and t he average roughness is about 40 nm over 600 m 600 m mirror surface T he reflectance is about 86.3% for the front mirror surface and 84.2% for the back mirror surface measured by using a 632 nm laser and an optical power meter The curvature of the mirror surface is mainly due to the residual stress of the SiO2 layer, and it can be further improved by removing the SiO2 layers on both sides of the mirror plate before Al deposition. The reflectance can be easily improved by a more uniform Al e -beam evaporation.

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96 Fig ure 4 6 SEM pict ures of fabricated 1 D hinged bimorph micro mirrors A ) A single -side coated micromirror with 300 m bimorph beam length, B) and C) a dual reflective micromirror with 200 m bimorph beam length. Inset s show the details of the bimorph/heater structure the b oth mirror surfaces with single -crystal silicon in between and wire -bonded device. Fig ure 4 7 Line scan s of the fabricated dual reflective mirror surfaces by a white light optical profilometer A) front -side mirror surface an d B) back -side mirror surface. 500 m SCS 100 m 5 0 m Front side mirror surface Backside mirror surface C 1 mm Pt Al/SiO2 SiO 2 20 m 20 m B 1mm A Mirror Bimorph beams Single crystal silicon(SCS) (~20m thick) Heater Substrate A B

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97 4.2.3 Steady State Equivalent Circuit The steady -state operation of the heater aside 1 D electrothermal micromirror can be modeled by a simplified equivalent circuit a s shown in Figure 4 8 The model re presents the components of the device both in thermal and electrical domains with its parameters defined in Table 4 1. Since the heat is generated at one end of the beam, the temperature distribution along the beam is approximately linear assuming the radiation and convection ha ve negligible effect. That is, the average temperature change on the bimorph beam can be expressed as: 2 / ) (l h bT T T (4 1 ) Fig ure 4 8. Equivalent circuit model of a 1 D hinged-bimorph and heater aside electrothermal micromirror hT is the highest temperature change o n the bimorph beam that is generated by the Joule heating of the heater wit h an applied voltage V [137] : 2 / ] 1 4 1 [0 2 E T hR R V T (4 2 ) and lT is t he lowest temperature change at the mirror end of the bimorph beams which is related to hT as : ) (br Tb br h lR R R T T (4 3 ) Thermal Domain Electrical Domain hTlT 0T V + h ET R 0 0 ER PtR 1 TR ETR V P2 TmR TbR CmR CbR 2 TR

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98 T able 4 1. Parameters used in the steady -state lumped element model for 1 -D micromirror Parameters Description Electrical domain V Voltage applied to the heater RE0 Electrical resistance of the heater at substrate temperature RET Electrical resistance of the heater at temperature Th Temperature coefficient of resistance of the heater Thermal domain RT1 Thermal resistance between heater and substrate RPt Thermal resistance of Pt connecting the heater and the pad P Power generated by the heater RTb Thermal resistance of the bimorph beams RT2 Thermal resistance between bimorph beams and the mirror plate RTm Ther mal resistance of the mirror plate RCb Thermal resistance of the bimorph beams due to convection RCm Thermal resistance of the mirror plate due to convection RT Thermal resistance of the whole device Th Maximum temperature on the heater T l Temperature on the bimorph beam at the mirror end And Rbr can be expressed by: )] //( //[2 cm Tm cm T cb brR R R R R R (4 4 ) Th erefore, the rotation angle can be related to the applied voltage as: ) 1 ( 4 1 4 10 2 br Tb br E T b b T TR R R R R V t L (4 5 ) Note that the total thermal resistance RT is si gnificantly affected by the structure of the device. RT could be very small i f there are Si and Al at both ends of the bimorph beams connecting to the substrate and the mirror plate, so the voltage required to heat the bimorph beams is large according to E quation 4 5 Plot (a) in Figure 4 9 shows the calculated results of the rotation angle versus the applied voltage in this case. However, if the device is designed to have pure oxide layer at both connecting ends, RT can be greatly increased due to the low thermal conductivity of the meshed thin oxide layer. In this case, the required voltage for the same rotation angle can be greatly reduced, which is shown in plot (b). By comparing the two plots,

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99 one can see the significant effect of the oxide thermal isol ation regions for reducing the driving voltage. Figure 4 9 ( c ) shows the experimental result from an actually fabricated device with partial Si etch under thermal isolation. Zoom in SEM pictures of the thermal isolation region with partial and complete Si undercut are shown in Figure 4 10. Fig ure 4 9 V oltage versus rotation angle from steady -state model calculations and an actually fabricated 1 D hinged-bimorph micromirror In the SEM picture shown in Figure 4 10 A), the SiO2 th ermal isolation region with partial Si undercut shows a shining ring around the meshed holes, other area with Si underneath shows gray color. This is due to the different etching rate on the meshed SiO2 regions and the bimorph beams caused by loading effec t in Si DRIE [159] The bimorph beam region has larger opening area than the meshed SiO2 region, and experienced faster Si etching rate. So the bimorph beams were released before the SiO2 was completely undercut Figure 4 10 B) shows a device by adding over etch after the bimorph beams were released. The SiO2 layer with complete Si undercut from another device exhibits some buckling and was broken at some points.

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100 Figure 4 10. SEM s of thermal isolation region A ) with partial Si undercut, and B) with complete Si undercut. 4.2.4 Experimental Results 4.2.4.1 Static r esponse The static responses of two fabricated devices with different bimorph lengths have been characterized by employing the experimental se tup used for characterizing scanning angle s of TTP micromirror s in s ection 3.3.2. Figure 4 1 1 shows the measured mirror rotation angles and heater resistances versus the applied voltages for both devices with the respective bimorph length s (300 (200 at applied voltages of 12.5 V (0.66 mW) and 17 V (0.69 mW) have been obtained. Good linear correlation (within 2.7% error from the linear fit ) of the applied voltages with both the ro tation angles and the heater resistances have been observed over large scanning ranges of about 80 and 60 respectively for the two devices. Buckled SiO 2 with complete Si undercut 50 m Pure SiO 2 SiO 2 with Si underneath A B

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101 Figure 4 1 1 Measured applied voltage versus rotation angle and heater electrical re sistance of 1 D hinged -bimorph micromirrors. A B) with 300long bimorph, and C -D) with 200long bimorph. Figure 4 1 2 Comparison of theoretical calculation FEM simulation and e xperimental results of 1 D hinged -bimorph micromirrors. A ) 300 long bimorph mirror, B) 200 long b imorph mirror. A B C D A B

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102 The temperature on the heater can be calculated using the measured TCR of Pt (0.0029/C) T herefore the average temperature on the bimorph beams can be estimated, which is about 87 % of the heater temperature from the calculation of the ste ady state model. Figure 4 1 2 shows the comparison between the experimental results and the FEM results and theoretical calculation for the two types of device s where the maximum error is about 11 %. 4.2.4.2 Dynamic r esponse and c ircumferential s can An experimental setup with two laser beams incident on both surfaces of the mirror plate was used to validate the concept of full -circumferential scanning, as shown in Figure 4 1 3 A ). A dual -reflective micromirror with a periment. To avoid block ing the reflected beam by the thick bulk Si of the device frame, the substrate at the front side of the mirror was diced off by a dicing saw T he device was then packaged on a thin glass slide. Circumferential scanning pattern was o bserved, but the scanning arc range decays significantly w hen the frequency of driving voltages increases. Figure 4 1 3 B) shows the scanning pattern under a 9 Vpp sinusoidal voltage at the devices resonant frequency about 428 Hz with a 4.5 V DC bias. The scanning arc angle is close to 90 from either mirror surface. This significant decay is believed to be due to the slow thermal response of the device. Fi gure 4 1 3 Circumferential scanning experiment A ) E xperimental setup, B) Circumferential scanning pattern generated at resonance. B A

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103 A position sensitive detector (PSD) was employed for the frequency response measurement A 0.5 -Vpp sinusoidal signal with a 2.5 -Vdc bias was applied to the device with the reflected laser beam from the micromirror incident on the PSD. By sweeping the frequency of the sinusoidal signal, the amplitude of the deflection angles at different frequencies is represented by the PSD output signal The measurement result is shown in Figure 4 1 4 The mechanica l resonant frequency is 428Hz. However, the thermal response of this device is very slow. A s shown in the plot the 3 dB cut -off frequency is only about 0.3Hz. Fi gure 4 1 4 Measured frequency response of a dual reflective 1 D mirror The reason of this s low thermal response is believed to be caused by the good thermal isolation of the meshed oxide as well as the large thermal capacitance of the bulk Si remaining underneath the heater region. According to the transient analysis p resented in section 3.4.2., the thermal cut -off frequency is inversely proportional to the thermal capacitance of the active heating region. As mentioned in previous sections, at the last release step of the device fabrication as shown in Figure 4 4 I), in complete Si isotropic etch under the meshed oxide regions have been observed for most devices. 428Hz 3 dB

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104 The 1 D hinged bimorph design presented above uses heater aside bimorph structure and increases the stable DC actuation range compared to previous heater along-bimorph structure. However, because the heater is very close to the device substrate the redundant Si underneath the heater cause s large thermal capacitance of the active heating region which results in a slow thermal response. New designs need to be deve loped either by increasing the DC actuation range o r improving the thermal response, or by combining both, to realize at least a 90 mechanical scan angle at above 30 Hz for real -time FCS imaging requirement 4 3 Dual -Folded -Bimorph Actuator and FCS Micro mirrors The previous section presents an FCS micromirror design with dual reflective surfaces. However, this design still suffers from several problems that prohibit its practical usage for FCS First, the slow thermal response of the heater aside -bimorph design prohibits the achievable scanning range at real -time imaging speed and only about half -circumferential scan range has been obtained. Second, the 1 D hinged bimorph structure has a rotation axis around the bimorph s substrate base resulting in an una voidable mirror center shift during the actuation For a 1 -mm mirror plate actuated by a 300 m long bimorph, the mirror center is displaced vertically over 900 m and laterally over 170 m for a 90 mechanical rotation. This mirror center shift significantly impairs the effective mirror aperture size and can easily cause optical misalignment duri ng scan. In addition, the undesired center shift is also proportional to the bimorph beam length which prohibits further improvement of the scan range. Last ly the previous design has a large initial tilting angle and a large initial height on the tip of t he mirror plate, resulting in a large space required for the device to operate. This makes the alignment and packaging of the imaging endoscope to be very complicated and limit the endoscope miniaturization Th is section presents a dual -folded bimorph ( DFB) actuator design that solves all the above addr essed problems. This new design enables a fast scanning with full -circumferential

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105 range at real time imaging speed, realizes a stationary rotation axis along the mirror center with an initially flat and un -elevated mirror rest position. 4.3.1 Dual -Folded -Bimorph Actuator Design Fig ure 4 1 5 A ) shows the 3 D model of the mirror design built in Intellisuite The mirror plate is suspended by two sets of DFB actuators Each set is composed of an outer and inne r actuator folded in series with equal lengths The other end of the DFB actuator is connected to the device substrate. Al and SiO2 are still used for t he bimorph materials with Pt heaters embedded in between. After the thin films are released, the bimorph beams of both outer and inner actuators curl upward due to the residual stress, but the mirror plate still remains at the original position with zero shift and rotation, as can be seen in the SEMs in Figure 4 1 7 Figure 4 1 5 3 D models of the dual reflective micromirror based on the DFB actuator design by IntelliSuite (Device Type I, substrate not shown ). A ) A t initial posit ion, B ) t he inner a ctuator being excited and C) the outer actuator being excited B C Inner actuator Mirror plate Rotation axis Oute r actuator A

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106 Wh en temperature change is introduced to the actuators by applying voltages to the heaters, both actuators tend to bend flat due to the different thermal expansions of the Al and SiO2 layers, but they rotate the mirror plate in the opposite directions, as shown in Figure 4 1 5 B) and C) Therefore, by alternatively driving the two actuators to provide 45 rotation angle, a 90 mechanical angle of the mirror rotation, i.e., a 180 optical scan angle, can be obtained from either mirror surface, resulting in a f ull -circumferential scanning by combining the two scanning optical beams. Compared to the previous hinged -bimorph design, this novel DFB actuator significantly increases the achievable rotation angle by combining the actuation of the two opposite actuators and it realizes a minimized mirror center shift by having the mirror rotation axis almost aligned to the mirror center axis. It is found by simple geometry analysis that the position for a minimum mirror center shift is to align the mirror center axis to the bimorph midpoint, and the mirror center shifts maximally 6% of the bimorph beam length, e.g., 45 mechanical rotation. Furthermore, the initially -flat mirror position greatly facilitates the assembly and optical alignment of the MEMS mirror in the imaging endoscope. 4.3. 2 Device Fabrication The device can be fabricate d using the same self aligned deep trench process as presented in section 4.2.2. A modified schematic of the process flow is illustrated in Fi gure 4 1 6 which starts from the completion of the front -side processes on top of the device layer of a n SOI wafer with multi layer deposition and etching steps that define the Al/ SiO2 bimorph beams and front side mirror surface The SiO2 layer on the mirror plate is etched before Al deposition to improve the mirror flatness, and the Al layer on the bimorph is extend ed to the device substrate to increase the thermal conductance for a faster thermal response. The addition of this Al metal

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107 layer can also reinforce the bimorphs anchoring strength to the device substrate and thus provide better device reliability The ba ckside process es start s from the SiO2 etch to define the backside mirror region and a thick layer of photoresist is used to protect the device substrate B) Si DRIE is then performed on the backside followed by an RIE to etch the exposed buried oxide layer C) After that, an Al layer is deposited for the reflective coating of the backside mirror surface D ) Then an isotropic Si etch is followed to remove the Si walls, and the SiO2 layer underneath is etched to separate the connection between the mirror plat e and the bimorph region E ). Finally, the device is flipped over and released by an anisotropic Si etch F ) and an isotropic Si etch G ). Fi gure 4 1 6 F abrication process flo w for a DFB -FCS micro mirror Fig ure 4 1 7 shows the SEM pictures of some fabricated devices The Type I device in A ) has two sets of actuators and the Type II in B ) has only one with an open mirror end The Type II design is to further avo id the beam blocking b y the actuators and substrate Note that the Si substrate at the mirrors open end can be removed after the last release step in Fig ure 4 1 6 G ). B C Si SiO 2 Pt Al P.R Bimorph beams Mirror plate A D E F G

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108 Both the Al and SiO2 layers are about 1 The bimorph beams are 300 0.8 mm2 with a device footprint of about 22 mm2. The quality of the mirror surface is measured by a Wyko white -light interferometer By remo ving the highly stresse d SiO2 layer on the mirror plate, an improved mirror flatness with the radius of curvature about 0.65 m and a surface roughness of 30 nm is obtained Fi gure 4 1 7 SEMs of fabricated DFB -FCS micromirror. A ) Type I and B) Type II. Insets showing backside mirror edge, Si frame and Al at substrate base. 4.3. 3 Lumped Element Model The device is operated using alternative excitation of the two bimorph actuators that are close ly located T herefore the thermal coupling effect of the two actuators need s to be considered L umped element model s with each actuator being individually excited are built to analyze the device s steady -state and transient behavior s under the thermal coup ling effect as A Mirror plate SCS Backside mirror (Al) 1 mm Inner actuator Outer actuator B Al

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109 shown in Figure 4 1 8 which are the schematics of the equivalent circuit s with the parameters listed in Table 4 2. Figure 4 1 8 The LEM equivalent circuit s of DFB -FCS micro mirror s. A) with the outer bimorph actua tor e xcited and B) the inner actuator excited T able 4 2. Parameters used in the lumped element model s for the DFB -FCS mirror Parameters Description Outer bimorph actuator P o T0 Tim Tio Toi Toh Tos Rm Cm Rai C ai Power generated by th e outer actuator Ambient temperature T emperature of inner actuator at mirror end T emperature of inner actuator at outer end T emperature of outer actuator at inner end Maximum t emperature on outer actuator T emperature on outer actuator at substrate end The rmal resistance through the mirror plate Thermal capacitance of the mirror plate Thermal resistance of the inner actuator Thermal capacitance of the inner actuator T oh T o s T 0 R aos R aoi R oi R ai T io T oi T im R caos R caoi R coi R cai R s R m C aoi C oi C ai C aos C s C m T o i T o s T 0 R ao R oi R aio R aim T ih T io T im R cao R coi R caio R caim R s R m C oi C aio C aim C ao C s C m R os R om R is R im B A P i P o

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110 Table 4 2. Continued Parameters Description R cai Roi Coi Rcoi Raoi Caoi Rcaoi Raos Caos Rcaos Rs Cs Ros R om Thermal resistance through convection on inner actuator Thermal resistance between inner and outer actuator Thermal capacitance between inner and outer actuator Thermal resistance through convectio n between inner and outer actuator Thermal resistance of the outer actuator between Toh and Toi Thermal capacitance of the outer actuator between Toh and Toi Thermal resistance through convection on the outer actuator between Toh and Toi Thermal resistance of the outer actuator between Toh and Tos Thermal capacitance of the outer actuator between Toh and Tos Thermal resistance through convection on the outer actuator between Toh and Tos Thermal resistance through the substrate Thermal capacitance through th e substrate Thermal resistance seen from the maximum temperature to the substrate Thermal resistance seen from the maximum temperature to the mirror Inner bimorp h actuator P i Tih Rai m Cai m Rcai m Ra i o Ca i o Rc a i o Rao Cao Rcao Ris R im Power generated by th e inner actuator Maximum t emperature on inner actuator Thermal resistance of the inner actuator between Tih and Tim Thermal capacitance of the inner actuator between Tih and Tim Thermal resistance through convection on inner actuator between Tih and Tim Th ermal resistance of the inner actuator between Tih and Tio Thermal capacitance of the inner actuator between Tih and Tio Thermal resistance through convection on inner actuator between Tih and Tio Thermal resistance of the outer actuator Thermal capacitan ce of the outer actuator Thermal resistance through convection on the outer actuator Thermal resistance seen from the maximum temperature to the substrate Thermal resistance seen from the maximum temperature to the mirror The model is lumped the same wa y as described in section 3.4.2 by splitting the excited actuator from its maximum temperature point, and the same derivation as in section 3.4.2 can be applied to obtain the maximum temperature change s for both inner and outer actuators with respect to th e time and a sinusoidal input: )] / exp( ) [exp( / 1 / ) (o o a o oht t j j C P t T (4 6) )] / exp( ) [exp( / 1 / ) (i i a i iht t j j C P t T (4 7)

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111 the long time solutions reduce to: o a t j o ohj C e P T / 1 / (4 -8) i a t j i ihj C e P T / 1 / (4 -9) The two actuators share the same thermal capacitance aC but they see different thermal resistances in the netw ork : om os oR R R // and im is iR R R // as shown in Figure 4 1 8 To improve the thermal response and device reliability the outer actuator is connected to the device substrate by an additional Al layer. But the inner actuator is still well thermally isolated from the mirror plate by pure SiO2 connection to minimize the heating on the mirror plate. Therefore, om osR R and i os oR R R A rough estimation on the fabricated device gives 8 20 K /W and 2200 K /W for outer and inner actuator respectively. The excitation of one actuator also induce s a temperature increase on the unexcited actuator as they are closely located: i a t j o oi oij C e P A T / 1 / (4 10) o a t j i io ij C e P A T / 1 /0 (4 11) oiA and ioA are coefficients with values between 0 to 1 denoting the portion of the heat distributed on the unexcited actuator T he two actuators are oppositely oriented and rotate the mirror plate in the opposite direction. Therefore, t he thermal coupling, i.e., the temperature increase on the unexcited actuator will decrease the effective achievable rotation of the mirror plate. The amplitude of the effective rotation angle by exciting each actuator can be expressed by:

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112 2 2) ( 1 / ) ( 1 /i i o oi o o o oi oh oR P A R P T T (4 12) 2 2) ( 1 / ) ( 1 /o o i io i i i io ih i R P A R P T T (4 13) Due to the different time constant ( i o ), the frequency response by individually exciting the two actuators behaves differently, as illustrated in Figure 4 1 9 based on the estimated thermal time constants It is interesting to see that the effective frequency response of the outer actuator has a broader bandwidth and an amplitude increase at low frequency before the cut off. This will eventually increase the bandwidth of the bi -directional scan by combining the two differentially excited actuators, which will be shown in the following subsection. Figure 4 1 9 F requency response s of the DFB actuator predicted by the LEM A) outer actuator exci ted, and B) inner actuator excited. 4.3. 4 Device Characterization A laser and a screen were used to characterize the rotation angles of the device By combining the bi -directional rotation from both inner and outer actuators, the two types of devices both demonstrated a maximum mechanical rotation of over 45 at less than 12 V driving voltages Fig ure 4 19 shows the measured results of a Type II device with the bi io ihT T ihT ioT oi ohT T ohT oiT B A

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113 directional rotation angles versus the actuation voltages for both the inner and outer actuators The response s at different actuation frequencies were also measured, and over 90 rotation angle was obtained at frequencies of up to 60 Hz As discussed in the previous subsection, the angular response of the inner and outer actuator is not symmetric due to their different thermal paths The outer actuator is closer to the substrate and thus has smaller thermal resistance which in turn results in smaller angular responsivity than that of the inner actuator, as shown in Figure 4 20. Figure 4 20. Measured DC and AC response of rotation angle versus voltage of a Type II DFB FCS micromirror. The devices frequency response is characterized and the results are close to the thermal coupling mode l prediction As shown in Figure 4 2 1 t he scan amplitude of the outer actuator increases at low frequency a nd reaches a peak at about 30 Hz, but the inner actuator behaves normally with a 3dB cut -off frequency at about 14 Hz Fig ure 4 20 also shows the ef fective frequency response of the amplitude by combining the bi -directional scan of both actuators obtained by different ially driving the two A smooth er response of the effective bi -directional

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114 scan is observed and the 3 dB cut -off frequency is above the devices mechanical resonance of 447 Hz Figure 4 2 1 Measured frequency response of a Type II DFB FCS micromirror. Fig ure 4 2 2 shows m icroscopic pictures of the devices at different actuation position s, in wh ich D ) shows the mirror being differentially actuated at 60 Hz, and a stationary rotation axis along the mirror center can be clearly observed Fig ure 4 2 3 shows the FCS patterns obtained respectively from the Type I and Type II devices by using two laser light sources incident on both surfaces of the mirror plate s Both devices were driven at 30 Hz with their inner and outer actuators being differentially actuated As shown in the pictures, the output beams from both the mirror surfaces overlap with each other, which demonstrate that the mirrors mechanical rotation angles are larger than 90 The Type I device still has the beam blocking problem and the output beams were partially blocked Second harmonic M echanical resonance

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115 by the device su bstrate as in A ), while the Type II device with the open mirror end design fully realizes the FCS function without any beam blocking as in B). Figure 4 2 2 Microscopic pictures of a Type II D FB FCS micromirror. A) A t initial rest position, B) outer actuator act uated, C) inner actuator actuated, D ) differentially actuated at 6 0 Hz showing a fixed rotation axis and E ) a Type I device with inner actuators actuated showing the ba ckside mirror surface. Figure 4 2 3 Full -circumferential scanning patterns at 30Hz by DFB FCS micromirrors A ) Type I device with optical beam partially blocked by device substrate and B) Type II device without beam blocking. E Backside mirror surface A B D Fixed rotation axis C B Dev ice A Beam blocked

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116 4 4 Light Weight Micromirrors As presented in previous chapters t he currently developed electrothermal micromirrors share the same bulk Si supported structure to maintain the mirror flatness for large aperture size. The S the mass of the bulk Si increases quadratically. T he large mass of the mirror plate significantly affect s the device performances in several aspects. The first is that it l imits the agility of micromirrors especially for those ones with large aperture size (>1 mm). For example, with a 1 mm by 1 mm sized mirror plate as described in s ection 4.2.4, the measured resonant frequency is 336 Hz By scaling the mirror size to 3 mm, the resonant frequency can be predicted to be below 100 Hz, so scanning speed for certain required scanning angle may drop significantly. Another effect is that the bulk Si induces large thermal capacitance on the mirror plate and the heat generated during the actuation takes a longer time to dissipate, which limits the devices thermal response speed In addition, the heat accumulated on the mirror plate induces thermal stress on the reflective coating layer which then degrades the surface flatness during actuation. An alternative option to increase the heat dissipation on the mirror plate is by increasing the heat transfer by convection effect, this can be realized by increasing the surface area of the mirror plate. Therefore, a process for fabricating th e mirror plate with reduced mass and increased surface area is required. Numerous fabrication methods have been developed to fabricate light -weight and optically flat micromirrors For instance, Kaiser et al. [160] used vertical nitride ribs to stiffen the mirror membrane and reduce the curvature of the mirror surface due to the internal stress in the nitride and the metal layer, which is accomplishe d by initially etching the sacrificial PSG and thermal oxide with hydrofluoric acid solution to form a mold that was filled with nitride to create the

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117 st iffening lattice supporting the mirror. Drake et al. [161] used a surface micromachining process and PSG sacrificial layer to create Poly Silicon ribs to support the mirror. These methods are based on relatively complicated process es that inv olve sacrificial structures As introduced in s ection 4.2.2, a self aligned deep trench process enables the backside mirror pattern to be defined in deep trench. By using a similar process, patterns can be defined on the backside of the mirror plate which enables a meshed Si rib structure with a significantly reduced mass and increased surface area. 4. 4 1 Self Aligned Deep Trench Process A novel self aligned deep trench fabrication process for light -weight micromirrors is proposed. This process is based on the backside patterning and self aligned deep trench etching of silicon substrate, which forms the mirror plate with meshed S i ribs to guarantee the flatness of the mirror surface and effectively reduce the mass of the mirror plate With the reduced mass o f the mirror plate and the greater stiffness, the resonant frequency of the mirror can be increased which expands the bandwidth of the device. At the same time, the ribbed structure significantly increases the surface area of the mirror plate so that the h eat transfer by convection on the mirror is improved. T ogether with a reduced thermal capacitance, the devices thermal response and the thermally induced mirror curvature can be improved. The complete process flow of the proposed fabrication process for a 1 D electrothermal micromirror using a n SOI wafer is shown in Figure 4 2 4 The front -side process for defining heaters, bimorphs and mirror structures is similar to the process for fabricating dual reflective mirrors as shown in Figure 4 4. Note that a Si O2 etch on the mirror plate is added to remove the compressive ly -stressed oxide layer for improving the mirror flatness.

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118 Figure 4 2 4 Self aligned deep trench process flow for fabricating light -weight micromirror Si SiO 2 Pt F ). Backside photoresist pattern ing E ). Backside alignment and SiO 2 etch I ). Partial b uried SiO 2 etch J ). Backside Si isotropic etch Al P.R. K ). Buried SiO 2 etch and Si etch L ). Remove P.R. and buried SiO 2 etch M ). Front side Si anistropic etch N ). Front side Si isotropic etch A ). Pt sputtering and lift off C ). Al e beam deposition and lift off D ). SiO 2 etch H ). Deep trench photoresist reflow G). Back side Si anisotropic etch B ). Dielectric SiO 2 and mirror SiO 2 etch

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119 The detailed description for the backside process flow from Figure 4 24 E ) is as flowing : E ): Backside alignment and SiO2 etch to define the meshed mirror rib pattern. F ): Patterning photoresist on backside for the DRIE mask on device substrate G ): Backside Si anisot ropic etch down to the buried oxide layer of the SOI wafer ; SiO2 on top of the ribs is consumed during the Si etch. The wafer is diced after this step and the following steps are performed at the die level. H ): Photoresist is dispensed into the backside c avity outside the mirror region area and a photoresist reflow baking p rocess is performed to protect this area during the following etching process. I): A partial SiO2 dry etch is performed on the buried oxide layer inside the deep trench of the meshed Si walls ; the etching is partially done about half way to generate a thickness difference between these deep trench area and the un -etched buried oxide area is protected by photoresist and Si walls. J ): A Si isotropic etch is performed to undercut the Si wall s so that the baskside of the mirror plate with a rib pattern formed by the thickness difference of the buried SiO2 in step I ) is generated. K ): Continue the SiO2 dry etch to completely remove the partially etched buried SiO2 in step I) so that the Si ins ide the meshed ribs is exposed, the buried SiO2 layer on the rib area where it has not been etched previously can be partially left as the mask for the following Si etch, therefore a ribbed Si structure under the mirror surface can be formed with a Si anis otropic etch reaching the Al layer. L ): A plasma etch process is then performed to remove the photoresist in the cavity, and the SiO2 dry etch is continued to remove the remaining buried oxide layer.

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120 M ) and N ): The device is then flipped over and a front -side anisotropic etch followed by isotropic undercut as in the release step for normal micromirrors will finally release the device. This self aligned deep trench process can be applied to all the electrothermal micromirror designs presented in this diss ertation 4. 4 2 Experimental Results LSF -LVD actuator based TTP mirrors have been fabricated with the l ight -weight structure using the above self aligned deep trench process as shown in Figure 4 25. Figure 4 2 5 Layout and SEM s of the l ight -weight micromirrors. A) Type I B) Type II light weight structure layout, C) SEM of a released Type II device and D) SEM of the backside of a Type I device before the last step of backside Si etch Inset SEM showing the l ight -weight pattern transferred on the buried SiO2 layer. D 2 mm B A SiO 2 Si C 2 mm

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121 Figure 4 2 5 illustrates the layout of the light -weight structures and SEM pictures of some fabricated devices. The mirror plate is 3 mm by 3 mm sup ported by 4 pairs of LSF LVD actuators at each side. The LSF -LVD actuators three bimorph len 400ight -weight structure with different meshing densit ies are fabricated. The mass es of the mirror plate s are r educed to 21.8% and 17.7% for Type I and Type II devices respectively Both devices demonstrate similar piston and tip tilt scan range to the device with complete Si layer underneath the mirror plate due to the same actua tor parameters and mirror size Figure 4 2 6 shows the DC response of the piston scan and tiptilt range versus the applied voltages of a Type I device. About 0.45-mm piston and 6.5 tip-tilt scan are obtained at voltages less than 4 V. Figure 4 2 6 DC response of the piston and tip-tilt scan versus the applied voltages for a Type II light -weight micromirror The device s mechanical resonances are characterized by applying a voltage of 2Vdc+0. 4Vac and sweeping the frequenc y of the ac signal. A PSD is used to track the scan range of the reflected beam from the mirror surface. Resonance frequencies of the piston and tip -tilt B A

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122 motion for a micromirror with a complete Si mirror plate and the two lig ht -weight micro mirrors are measured and compared in Table 4 2. The resonance frequency of the piston motion can also be predicted as a simple spring -mass system, given by: M K f2 1 (4 1 4 ) As shown in Table 4 3 the piston resonance frequencies of the light -weight mirrors are reasonably close to the calculation results by Equation 4 1 4 compared with the measured result of the micromirror with a complete Si mirror plate The mirror surface quality is measured by a Wyko white light interferometer D egradation of the surface quality has been observed for the l ight -weight mirrors compared to the complete Si mirror plate For a Type II device, the curvature radius was decreased from 0.80 m to about 0.15 m the average roughness increased from 35 nm to be about 80 nm and the reflectivity from about 96 % to 62 %. This degradation is developed during the high temp erature processes of DRIE and RIE steps, in which some permanent deformation on the Al thin film occurred. Figure 4 2 7 Microscopic images of the Type II device mirror surface with piston actuation at differen t voltages A) 0 V B) 1 V and C) 2.5 V. Further degradation wa s even observed during device operation As shown in Figure 4 2 7 the electrothermal actuation causes temperature increases on the mirror plate and the mirror surface corrugates due to the ther mal expansion of the Al thin film. To improve the mirror B A C

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123 surface quality, a multilayer thin film structure with a careful thickness design, as presented in section 3.4, is needed for the thermal stress compensation. Table 4 3 Comparison of resonance frequ encies of complete -Si device and l ight -weight devices Device Type Mirror plate mass Piston resonance frequency (Hz) Tip tilt r esonance frequency (Hz) Calculated Measured Complete Si 100 % 242 357 L ight weight I 21.8 % 528 419 445 L ight weight II 17.7 % 575 542 560 4 5 Summary This chapter proposes a novel full circumferential scanning endoscope design by using a large angle dual reflective micromirror A novel process has been developed for fabricating the dual -reflective micromirror and the feasibility has been verified on a hinged bimorph 1 D micromirror The fabricated device demonstrated rotation angles larger than 90 at steady state and about half circumferential scanning range has been obtained wh en operated at resonance. To overcome the drawbacks of the hinged-bimorph design, a novel dual -folded-bimorph (DFB) actuator is designed and DFB -FCS mirrors are fabricated. The new devices solve the mirror center shift and large initial tilting problems, a nd can realize full -circumferential scan at real time imaging speed above 60 Hz. A novel self aligned deep trench process is also proposed to fabricate light-weight mirror plate s for large aperture and agile MEMS mirrors TTP mirrors with a large aperture size of 3 mm are fabricated using this new process and the devices demonstrate a significant improvement on the mechanical resonance frequencies due to the reduction of the mirror mass.

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124 CHAPTER 5 APPLICATIONS OF MEMS OPTICAL SCANNERS As discussed in Chapter 1, one of the main research goal s of this project is to develop miniaturized endoscopes that enable in vivo detection and diagnosis of precancerous lesions using advanced optical imaging techniques such as optical coherence tomography (OCT) and nonline ar optical (NLO) microscopy As the key technique for the miniatu rization of scanning endoscopes, several electrothermal MEMS micro -scanners dedicate d on improving scanning range and fill factor, have been presented in the previous chapters As a major ap plication of the developed MEMS optical scanners, the first part of th is chapter is focuse d on the development of MEMS -based endoscop ic probes including the probe design MEMS packaging and probe assembling Then OCT and NLO imaging results using the MEMS -based probes are presented. In addition, scanning microlens es with glass lenses mounted on a modified LSF -LVD actuator ha ve been developed and been used in confocal microscopy and optical coherence microscopy. Imaging results will be presented. The advant ages of large actuation range and actuation force of these novel electrothermal MEMS optical scanners also lend themselves to many other applications besides the biomedical imaging. In this research effort the great potentials of the developed MEMS optica l scanners in other applications such as optical phased array, free -space optical communications and especially a great opportunity for the miniaturization of the Fourier transform spectrometer s (FT S ) are investigated These applications are discussed in t he last section s of this chapter. 5. 1 MEMS -based Nonlinear Optical Endoscope The development of an MEMS based miniature endoscope for nonlinear optical imaging is collaborated with Dr. Dru Morrish and Prof. Min Gu in the Centre for Micro -Photonics, Swinbu rne University of Technology, Hawthorn, Australia

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125 5.1.1 Nonlinear Optical Imaging System Non linear optical imaging use s the nonlinear optical effects of tissue for imaging with high resolution up to sub-micron both axial ly and lateral ly NLO imaging ha s great potential to achieve in vivo diagnosis of cancers at a very early stage Among the nonlinear optical effects, t wo photon excitation fluorescence (TPEF) and second harmonic generation (SHG) currently attract enormous research interests in biomedical imaging applications since the introduction of femtosecond pulsed laser s in the near infrared wavelength range for greatly localized excitation [26 29] T he nonlinear optical imaging system developed at Swinburne University of Technology is capable of excitation and detection of both TPEF and SHG signal s [14, 53] Figure 5 1 Nonlinear optical imaging system. A ) Schematic of imaging setup, B) SEM picture of a double -clad photonic -crystal -fiber (Imag ed by L. Fu, with permission) Ti: Sapphire Laser Prechirp Unit Dichroic Mirror 40X Objective lens PMT Iris Filter Bandpass Filter Double -clad Photonic crystal fiber Transverse scanner Tissue sample Ti: Sapphire Laser Prechirp Unit Dichroic Mirror 40X Objective lens PMT Iris Filter Bandpass Filter Double -clad Photonic crystal fiber Transverse scanner Tissue sample A B

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126 A schematic of the imaging set up is shown in Figure 5 1 A ). A Ti: Sapphire laser (Spectra Physics Mai Tai) is used as the excitation light source which generates a 800 nm about 80 fs pulses at a repetition rate of 80 MHz w ith an output power of 850 mW. The laser pulses are deflected by a dichro ic mirror which is used to separate the excitation laser beam and the emission signal, then coupled into a double -clad photonic -crystal -fiber (DC PCF) by a micro cope objective lens (4 0x /0.65 NA). The DC PCF (Crystal Fiber A/S) is specifically designed for simultaneously delivering the excitation light and collecting the fluorescence signal. As shown in Figure 5 1 B), it has a n inner core diameter of 20 m which is used for single -mode propagation of the 800 nm near -infrared excitation light, and an inner cladding with diameter of 165 m and an outer diameter of 550 m used for the multimode propagation of the visible emission light. At the end of the fiber, a scanning mechanism is used to scan the excitation laser beam which is then focused by a 0.2 pitch GRIN lens onto the tissue surface. The emission light from the tissue is then collected by the fiber and the same focusing optics, and finally detected by a photomultiplier tube (PMT). Both TPEF and SHG signals can be excited To differentiate them a bandpass filter is inserted before the PMT. Both 1 D and 2 D MEMS electrothermal micromirrors developed by Jain et al. in [65, 115] have been used as the scanning mechanism in the imaging system L ine scan profiles and 3 D reconstructions of ex vivo tissue samples by the bench top system have been demonstrated with 1.5-m lateral resolution and 100 m penetration depth [53, 162] For further in vivo imaging of internal organs, miniaturized endoscopes must be developed. A s discussed in s ection 2.4, the previously used 2 D MEMS micromirror (as in Figure 2 4 B )) is limited by its effective scanning range due to the large shift of the mirror center and small aperture size. In addi tion the relatively low fill f actor of the mirror limits the miniaturization of the endoscope.

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127 5.1.2 Nonlinear Optical Endoscope Design and Imaging Results To solve the problems of the previously used 2 D micromirror, a TTP MEMS mirror based on the LSF -LVD actuator design (Type II) as described in Chapter 3 with a mirror size of 1 mm is used for the development of the nonlinear optical endoscope. A 3 D model of the endoscopic probe is shown in Figure 5 2. Figure 5 2 3 D model of the nonlin ear optical imaging endoscope A ) Assembly of the complete probe, B) probe body and C) probe cap. The probe consists of two parts : a probe body and a cap. T he two parts are designed in such a way that the probe body can be slide into the cap and form a se aled cavity for the MEMS mirror to perform 2 D scans. This also allows easy access for MEMS mirror integration, alignment and replacement A 45 slope is machined at one end of the probe body and a 0.5 mm deep pocket is formed for mounting the MEMS mirror. This pre -defined pocket is used for aligning the MEMS mirror. The space taken by the device substrate is saved by hiding it inside the probe. Five copper posts are threaded through the probe body and polished with a 45 slope A Fixed mirror GRIN lens C Fiber hole Copper post MEMS mirror B

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128 surface for the wire bonding of the MEMS mirror. It also has a hole at the lower part for delivering the optical fiber into the probe. The cap is machined with a 4.4 -mm diameter window at the forward end for fixing the GRIN lens, and a 1 mm by 1 mm slot is machined for fixing a 45 fi xed mirror inside The 45 fixed mirror is a prism coated with a metal film (Al). The excitation laser beam is delivered into the probe by the fiber, and deflected to the MEMS mirror by the fixed mirror The MEMS mirror surface is initially flat so that it is parallel to the fixed mirror. The reflected beam from the MEMS mirror is then focused by the GRIN lens onto the tissue surface. With a 2 D scan of the MEMS mirror, the tissue sample can be imaged at different depths so that a 3 D visualization can be o btained. The endoscope is 5 mm in diameter and 23 mm long. The GRIN lens is 4.4 mm by 1.8 mm and the optical path length from the fiber end to the GRIN lens is about 7.4 mm. The MEMS mirror has a significantly improved fill factor of 25 % with an aperture size of 1 mm on a 2 mm device footprint, compared to a 4.8% of the previous design (0.5 mm aperture size on a 2.7 1 .9 mm device footprint ). The new MEMS mirror enabl es a more compact endoscope design with an increased mirror aperture size. Figure 5 3 Pictures of fabricated NLO probe A -C) probe body, and D ) with MEMS mirror assembled into the probe. Copper posts bonded by silver epoxy MEMS mirror D 5 mm 23 mm Pocket for MEMS mirror Copper posts C B A

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129 Pictures of fabricated probe and the wire bonded MEMS mirror are shown in Figure 5 3. The endoscopic probe was custom -machined by TMR Engineering, Micanopy, FL. Lucite was chosen as the probe material as it is bio -compatible, offers natural dielectric for the embedded copper wires, and also has proper hardness for fine machining. A n MEMS mirror was mounted into the pre -machined pocke t on the 45 slope, and wire bonded to the copper posts. Silver epoxy was used for soldering the gold wires to the copper posts. Figure 5 4 Experimental results of the NLO probe with MEMS mirror assembled. A ) m easured optical scanning angles at different voltages and frequncies and B) p ictures of 2 D scanning patterns generated by the probe Before assembling the fiber and GRIN lens, the optical alignment of the endoscope was verified by an experimental set up with a laser be am passing through the hole that is used to hold the DC PCF fiber The scanning angle of the output beam from the GRIN lens window was then characterized by differential ac voltages with a dc bias applied to the two opposite actuators in both directions. A dc bias of about 3.2 V wa s needed on the four actuators for aligning the MEMS mirror to the center axis of the GRIN lens window. As the required scanning frequency for nonlinear imaging setup wa s around 10 Hz (7 lines/s), the optical scanning angles at 10 Hz and 20 Hz in both vertical and horizontal directions were g enerated at a DC bias of 3.2 V with B A B locked by GRIN lens window beyond 2V ac

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130 varying ac voltage amplitudes. Figure 5 4 (a) shows the testing results with 2D scanning patterns in (b). Figure 5 5 T wo photon fluorescence image stacks of fluorescent beads, scale bars 10 m. Figure 5 5 shows t ypical two -photon fluorescence images of fluorescent beads obtained by the endoscope The mirror plate was translated by a total of approximately 0.175 mm in the z direction. Each of the eight z sections is separated by approximately 0.025 mm. The axial sections are translated with respect to each other due to a small angle between the optical axis and the z axis translation of the MEMS mirror. The scanning range of the 3 -D (TTP ) MEMS mirror is sufficient for the acquisition of image stacks The endoscopic probe containing the 3 D scanning MEMS is of sufficiently small dimensions making it suitable for in vivo operation. 5. 2 MEMS -based Optical Coherence Tomography En doscope As discussed in Chapter 1, OCT is one of the most advanced and powerful biomedical imaging techniques that have been recently developed. Early endoscopic OCT (EOCT) probes using micromotors [47] can only perform side view imaging and has scan synchronization problem. MEMS m icromirrors are more versatile, co st -effective and thus more suitable to rea lize miniature EOCT probes Pan et al. reported the first MEMS -based endoscopic OCT using an electrothermal micromirror in 2001 [163] After that, various MEMS micromirrors have been 1 2 3 4 6 5 8 7

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131 developed, among which electrostatic micromirrors attract most interest due to the advantages of fast speed and low power consumption [58, 164] However, they need high driving voltages, which are typically on the order of 100 V; posing a potential risk to the patient. Electromagnetic micromirror -based OCT probes have also been investigated [165] but is limited by the complex assembling process and further miniaturization due to the permanent magnet required. Another common problem for electrostatic and electromagnetic micromirrors is the small fill factors that are typically less than 10%. One solution for the small fill factor problem is to use hidden actuators [91] but the processes for the formation of the mirror tops are quite complex and not reliable. Another s olution is to use electrothermal bimorph actuators [73, 112, 166, 167] As presented in this research effort, in addition to the advantages of large scan range (above 60) at low driving voltage (less than 10 V), the compact bimorph actuator enables relatively high fill factor up to 25% for a 3 -D micromirrors. These features make the thermal bimorph -based MEMS mirrors very suitable for miniaturizing OCT probes. This section discusses the research efforts on the EOCT probe development in the Biophotonics and Microsystems Laboratory (BML) at University of Florida. T he LSF LVD based 3 D MEMS mirrors presented in Chapter 3 are used, a 5.8 -mm diameter probe is developed, and 3 D in vivo images of a mou se tongue and ear have been successfully obtained with this MEMS -based OCT probe. This work is done with the OCT team at BML Dr Shuguang Guo, Jinjing Sun and Lin Liu. 5. 2 1 OCT System Figure 5 6 shows t he schematic of the time -domain OCT system The br oadband light source (DenseLight, DL BX9 CS3159A) has a minimum power of 15mW, a center wavelength of 1310 nm and FWHM of 75 nm which results in a 10 m axial resolution in air. The input light goes through the circulator and is divided by the 22 beam sp litter into the reference arm

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132 and the sample arm. The depth scanning at the reference arm is realized by a galvanometer based rapid scanning optical delay line (RSOD). The scanning rate of the RSOD is 1 kHz, and the carrier frequency of the OCT signal is 500 kHz, which is generated by the lateral shift of the rapid scanning mirror in the delay line. The transverse image scanning is realized by the MEMS mirror inside the endoscopic probe. The interference signal is then detected by a balanced photodetector, captured by a data acquisition card, and processed by a computer. Figure 5 6 Schematic of the time -domain OCT system 5. 2 2 OCT Endoscope Design and Imaging R esults A one piece probe as shown in Figure 5 7 is designed for easy optical alignment and the convenience of providing electrical connection to the MEMS mirror. The light from the sample arm propagates through a single -mode fiber (Corning SMF 28). T he jacket of the distal tip of the optical fiber is strip ped off and the fi ber is cut with an 8 degree angle at the end with an angled fiber cleaver (Cleaver MAX CI 08) to avoid 4% back reflection. A mechanical stop is used to fix the position of the fiber. The fiber is then connected to a GRIN lens with an optical UV glue to red uce the transmission loss caused by the index mismatch. The GRIN lens has a diameter of 700 m and a length of 1.95 mm; it is used to focus the optical beam on the sample. The MEMS mirror is used to realize transversal scanning. Gold wires are first bonded to the bonding pads on Light source 22 Beam splitter Circulator RSOD Photo detector Endoscopic probe DAQ PC

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133 the MEMS chip, and then the MEMS chip with the gold wires are assembled into the square pocket on the probe. The gold wires are connected to a small PCB board, which is connected to longer copper wires that come out of the probe from the backside slot, and connect to the control signals After probe assembly is completed, the entire probe is inserted into a flexible, biocompatible, transparent f luorinated e thylene p ropylene ( FEP) tube. To make the probe rigid, a wood stick with a lon g slot for wires is inserted into the FEP tube as well. The top and bottom of probe are both sealed with bio compatible glues After packaging, the outer diameter of the MEMS OCT probe is 5.8 mm. Figure 5 7 3 D model and pictures of the OCT probe A ) 3 -D model of the probe design, B) packaged 3 D MEMS OCT probe, C) c lose up of the assembled MEMS OCT probe. The packaged probe has been used for in vivo imaging experiments. A female athymic (nu/nu) nude mouse with a body weight of 20 g to 26 g (Harlan Laboratories, Indianapolis, IN) was used in the experiment. The mouse was anesthetized by injecting ketamine (100 mg/kg) and xylazine (10 mg/kg) intraperitoneally. Both 2 D and 3 D images of mouse tongue and mouse ear have been obtain ed, as shown in Fig ure 5 8 and Figure 5 9 respectively. The 2 D images are obtained at 2.5 frames per second. The two MEMS actuators on the circumferential direction are MEMS mirror A B C

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134 used for fast scanning, driven with 0~4V differential ramp voltage. And the other two MEMS actuators are used for slow scanning, driven with 0.5~3.5 V differential ramp voltage. This results in a 2.3 2.3 mm2 transverse area. Combined with the axial scanning of the RSOD, the volume of each 3 D image is 2.3 2.3 0 .6 mm3. The 2 D OCT imag e shown in Figure 5 8 A ) was acquired using the MEMS OCT probe at the base of mouse tongue. A stratified squamous keratinized epithelium (SSKE) and basement membrane (BM) corresponding to relatively strong signal band from the upper layers of the tissue st ructure were observed. Figure 5 8 2 D and 3 D in vivo OCT images of mouse tongue. A ) 2 D image of a mouse tongue. B) 3 D image of a mouse tongue. A 2 D OCT image of the mouse ear is shown in Figure 5 9 A). The ear thickness i s about 500 m. The dermal structures and the subcutaneously adjacent layers were observed using the MEMS OCT probe in this experiment. The cartilage (C) of the ear was represented in the middle by the dark band and the dense conjunctive capsule (cc) with two bright layers put around the cartilage. The superficial epidermis (E) of the mouse ear was detected at the first and last of the mouse ear longitudinally. The lower and upper areas from the dense conjunctive capsule to epidermis, the dermis (D), were o bserved Figure 5 9 B) shows the 3 D reconstruction image of SSKE BM 1cm A B

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135 the ear. The results show that using MEMS based OCT probe is very promising in biomedical study. Figure 5 9 In vivo images of mouse ear A ) 2 D image of a mouse ear B ) 3 -D image of a mouse ear 5.3 MEMS Microlens Scanners As discussed in Chapter 1, biomedical imaging techniques with cross -sectional imaging capabilities, such as optical coherence microscopy and confocal microscopy, require tunable microlenses capable of axial scan of the focal plane. Tunable ranges in order of 1 mm are desired for imaging deep into tissues. Liquid lenses with tunable focal length [122, 168] require large actuation voltages (>100V) and have slow response time (in seconds). Physically displacing microlenses by MEMS actuat ors are also investigated [54, 61] Particularly, electrothermal actuation has advantages of much larger actuation range and smaller driving voltage than other methods such as electrostatic actuation in [61] For instance, a large-vertical -displacement (LVD) electrothermal bimorph actuator with a 0.7 -mm actuati on range at 23 V was demonstrated [54] The novel lateral -shift free (LS F) LVD actuator design presented in Chapter 3 solves the large lateral shift and tilting problems in [54] However, the previous designs use polymer lenses formed by reflow at high temperature. Such polymer lenses have short lifetime for electrothermal B A E D D E C cc cc 1cm

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136 actuation. This section presents tunable microlens es with glass lens es assem bled on a modified, more robust LSF LVD actuator. 5.3. 1 New LSF LVD Lens Scanner Design and Fabrication The LSF LVD based lens scanner presented in Section 3. 3 has some small tilting (~ 1 ) of the mirror plate during the entire vertical scanning range This tilting mainly comes from the fabrication variations that cause mismatches among the four actuators structural parameters, including the thickness and width of the Pt heaters and bimorph beams This small tilting during the vertical actuation will cause aberration and is undesired for biomedical imaging Figure 5 10. L ens holder by modified LSF -LVD actuator design. A) Top view layout, and B) a simplified 3 D model built in Intellisuite. A new design of an LSF -LVD based lens s canner is proposed to further minimize this tilting effect in the current design. As shown in Figure 510, instead of a four actuator supported and point anchored structure, the new design combines two sets of actuators at each side of the lens holder and uses a line anchored bimorph along the lens holder edges. This is for the purpose B A LSF LVD actuator

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137 of better actuation stability with an easier control to minimize the residual lateral shift and tilting. The bimorph beams and frames are widened to gain a stiffer actuation for glass lenses A combined surface / bulk -micromachining process similar to the process in section 3.3 is used for device fabrication The BK7 glass microlenses are assembled to the MEMS lens holders by a UV -cured optical adhesive after the lens holders ar e released Fig ure 5 -1 1 shows the SEM pictures of fabricated devices including an unloaded lens holder and the lens holders with two types of glass lenses assembled The device footprint is 3 by 3 mm and the initial elevation of the unloaded lens holder is about 1 mm No obvious decrease of the elevation height was observed after both lenses were assembled Table 5 1 lists the glass lens parameters Figure 5 1 1 SEMs of fabricated lens scanners. A ) LSF LVD actuator, B) unloaded lens holder, C) assembled with lens A and D ) lens B A Al/ Pt/ SiO 2 Frame B 1 mm C 1 mm D 2 mm Assembled Glass Lenses

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138 Table 5 1 Glass l ens parameters (1300nm wavelength) Type Diameter (mm) Focal Length (mm) NA Beam Width Depth of Focus Mass (mg) A 0.7 2 0.26 4.7 27 0.29 B 1 3 0.25 5.0 30 1.97 5.3.2 Exp erimental Results Fig ure 5 1 2 shows the measured DC response of the vertical displacements versus the applied voltages on the two actuators at both sides Large displacements above 0.88 mm were obtained for both lenses at 3.75 V (~495 mW) Fig ure 5 1 3 show s the pictures of a lens B scanner being actuated at different voltages, and the microscopic images of a stack of transparency masks formed by the microlens using the same optical setup as presented in section 3.3 Each layer of the three stacked transpare ncy masks is 0.28 -mm thick with the top layer being placed about 4.8 mm under the microlens By actuating the lens scanner at different voltage levels, the three layers at different depths can be imaged Figure 5 1 2 DC response of vertical displacement v ersus voltag e

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139 Figure 5 1 3 Pictures of lens B scanner and microscopic images of transparency masks formed by the lens at different actuation voltages A ) 0V, B) 2 V, and C) 3.5 V. Figure 5 1 4 Frequency responses of the unloaded lens holder, the lens scanners assembled with lens A and lens B Fig ure 5 1 4 shows the measured frequency responses under voltages of 2 Vdc 0.4 Vac, the resonance frequencies of an u n loaded lens h older, a lens A scanner and lens B scanner are X Y B C A

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140 257 Hz, 158 Hz, and 79 Hz, respectively. There are small residual lateral shift and tilting during vertical actuation mainly due to some small differences among the bimorph actuators caused by process variatio ns with the same actuation voltage applied to both actuators for a lens B scanner An optimized voltage ratio (OVR) for the lateral shift and tilting miniaturization can be experimentally obtained by using a position-sensitive detector to track the shift of the light beam that is refracted by the microlens The lateral shift and tilting can be further minimized to below 7 m and 0.38 for the entire 0.88 mm vertical actuation range Fi g ure 5 1 5 shows the measured tilting angles in both directions with and without the OVR. Figure 5 1 5 T ilting angles of a lens B scanner with and without the OVR 5. 4 MEMS Optical Phased Arrays O ptical phased array (O PA ) techniq ues have demonstrated a wide range of applications including spatial light modulation, wave -front control/shaping, beam steering/aiming and micro

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141 displays [169] Particularly, optical phased steering that enables revolu tionary systems with random access pointing in free -space laser communication attracts the most research interests [170, 171] A variety of technical approaches for optical phased steering have been extensively investigated [71, 169, 172177] The conventional approach is by liquid crystal technology, which usually has small aperture size, small steering angle and dispersion problems and needs to be combined with large angle step -steering approaches [169] Other alternative methods include electrowetting [175] microlens array [176] vertical continuous OPA [177] and MEMS mirrors among which MEMS mirrors are dispersion-free and polarization invariant and have emerged as a promising s teering option with a possible simultaneous phase control [7 1, 174] Early developed MEMS mirror arrays such as the Digital Light Processor (DLP) by Texas Instruments [75] and mirror arrays by Lucent for optical communications [178] are limited by the small steering range without phase control, and are based on thin -film structures with small segmented apertures. Therefore, large OPAs comprising a large amount of these small segmented mirrors are impractical due to the high complexity of mirror control electronics. TTP micromirrors based on single -crystal -silicon structures have also b een demonstrated both by electrostatic and electrothermal actuations [65, 87, 88, 90, 91, 179, 180] Electrostatic actuation is faster and consumes low power consumption, but is limited by its small actuation range and large actuation voltage. For example, a 44 TTP mirror array by electrostatic comb drives was presented by Milanoic et al. and abou t 10 optical angle and obtained at with mirror sub aperture size up to 0.8 mm [90] Jung et al. also demonstrated a 33 electro aperture size and tip tilt angles less than 2 and piston deflection of 70 nm at 160 V [179] Another challenge for MEMS OPAs is the high fill factor. Electrostatic actuators usually occupy large area comparable to or even larger than mirror

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142 aperture size T herefore, dedicated wafer -bonding and assembly processes need to be employed to realize the high fill -factor as in [90] and [179] On the other hand, as demonstrated in the previous chapters, electrothermal actuation provides much larger actuation ran ge for large sub aperture size s in mm scale at small driving voltage s with much more compact actuator structure s, so electrothermal actuation has potential to realize high -fill -factor large arrays by relatively simple and high yield surface micromachining This section presents 44 electrothermal TTP mirror array s by the novel compact -folded, LSFLVD actuator. A phase control experiment by using two adjacent mirrors on the mirror array is also presented for the optical phased steering proof -of -concept. 5. 4.1 Mirror Array Design and Fabrication The TTP mirror array uses the same LSF LVD actuator design as in section 3.2. As shown in Figure 5 1 6 t he rectangular mirror aperture is supported by four identical actuators on the mirror edges. The piston displace ment is obtained by simultaneously driving the four actuators, while the tip tilt scan in both orthogonal directions can be generated by alternatively driving the two opposite actuators in each direction. A similar surface and bulk combined micromachining process as in section 3.2 can be used for the device fabrication. The process starts from surface micromachining of multilayer thin film deposition and etching steps to form the bimorphs, frames pattern and reflective coating of the mirror region. SOI wafe thick device layer is used to form the rigid frames and mirror plates. A Si DRIE is then performed from the backside to etch the bulk Si of SOI handling layer, the buried oxide layer works as an etch stop layer in this step so t hat the uniformity of the Si underneath the frames and mirror plates can be guaranteed. The devices are then finally released from the front -side by an anisotropic and an isotopic Si dry etch steps. In the last isotropic Si etch step, the width

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143 difference of the bimorphs and frames allows the majority of Si to be remained underneath the frames when a complete Si undercut of the bimorphs has been reached. Figure 5 1 6 SEM pictures of TTP mirror arrays. A ) D evice Type A with 31% fill -factor, and B) Type B with 65% fill -factor. The insets show one subaperture and the details of an LSF -LVD actuator. Fig ure 5 1 6 shows the SEM pictures of two types of devices w ith different subaperture size and fill -factor T he structure parameters summarized in Table 5 2 By increasing the subaperture size and decreasing the spacing between adjacent mirrors that is occupied by the actuators and the electrical wiring, the fill -factor is increased from 31% (Type A) to 65% (Type B). The mirror surface quality and fabrication uniformity of a Type B device is characterized by a Wyko NT9800 white light interferometer. The mirrors radius of curvature is about 0.68 m with an average roughness of about 40 nm as shown in Fig ure5 1 7 C) The average initial 2 m m 2 m m 2 m m Bimorph I Bimorph II Bimorph III Frame I Frame II A B

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144 el eva tion of all the 16 sub. Figure 5 1 7 Mirror surface measurement of a Type B device by a Wy ko NT9800 white light interferometer A ) T he entire array, B ) single mirror, and C) mirror surface deformation and roughness by a line scan. Table 5 2 Parameters of the fabricated mirror arrays Type Bimorph length ( ) Bimorph width ( ) Frame length ( ) Frame w idth Sub aperture size Fill factor A 100,200,100 40 200 200 62 500 31% B 100,200,100 28 200 200 44 9 0 0 65% 5.4.2 Device Characterization The static response of the piston displacement by applying the same voltage on all the four a ctuators was measured by an Olympus BX51 microscope and a QC200 geometry measuring system As shown in Fig ure 5 1 8 from both devices at less than 5 Vdc. Fig ure 5 1 8 A ) shows a good repeatability betwe en the forward and backward -driving of a device Type A after the burn in process. The same device was also operated for 10 times of forward -driving and an average 1.3% difference of the measured piston displacements at same voltages was observed, which is believed to be due to the measurement errors and ambient temperature fluctuations. There are small residual lateral shift A A A B A A line scan C

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145 and tilting of the mirror plate during the piston actuation As shown in Fig ure 5 1 8 B), a Type B device was characterized with a maxi mum 4.2 lateral shift and tilting for the entire 202 piston actuation. The lateral shift and tilting are mainly due to the variations of the actuator structures devel oped in the fabrication process A more uniform fabrication process can be employed to further minimize the lateral shift and tilting. Figure 5 1 8 Piston actuation DC response of the TTP mirror arrays. A ) P iston displacement versus the applied voltage of both devices; B) lateral shift and tilting angle versu s applied voltage of a Type B device. Figure 5 1 9 Tip -tilt actuation DC response of the TTP mirror arrays. A) Device Type A and B) Type B. A B A B

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146 The tip -tilt scan by exciting individual actuator are measured using an experimental setup with a He Ne laser incident on the MEMS mirror and a screen to track the pointing position of the reflected beam. Fig ure 5 1 9 shows the measured results for both devices Maximum tiptilt angles (optical angle) of about 64 (32) and 36 (18) in both orthogonal directions are obtained by the Type A and Type B device respectively. The tip tile angle is generated by the elevation difference of opposite actuators, and thus is directly related to the achievable piston displacement of the actuator and t he distance between the two opposite actuators, i.e. the mir ror aper t ure size. T he frequency .response s of both piston and tiptilt mode have been characterized The same experimental setup for the DC response measurement was used to measure the tip tilt frequency response, and a position-sensitive detector (PSD) was employed for the piston motion measurement due to the small deflection of the reflected beam by the piston motion. The resonance frequencies of the piston and tiptilt scan modes are 1.334 KH z, 3.147 KHz, and 785 Hz, 1.185 KHz for Type A and Type B devices respectively Fig ure 5 20 shows the measured frequency response of a Type B device at a voltage of 2Vdc+0.5Vac on all the four actuators for piston mode as in A ), and a same voltage with 180 phase difference of the ac signal on two opposite actuators for the tip-tilt scan mode as in B) Both plots in Figure 5 -20 show that the scan range drops at low frequency before the mechanical resonance due to the decay of thermal actuation, and a 3 -dB cut -off frequency about 42 Hz has been observed for the Type B device The thermal decay is also presented by the transient response of the tip tilt scan with single actuator excited As shown in Fig ure 5 2 1 both devices show an under -damped mechanical response combined with a thermal delay, the thermal response time (90% rising) for the Type A and Type B devices are about 30 ms and 12 ms respectively The faster thermal response of Type

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147 B device is due to its smaller thermal time constant to which small er thermal capacitance and resistance both contribute The actuator s smaller thermal capacitance is due to the less amount of Si underneath the narrower actuator frames, and the smaller thermal resistance is obtained by the actuator s better thermal conta ct to the device substrate with the additional Al metal layer at the junction. Figure 5 20. Frequency response s of Type B TTP mirror array. A ) P iston actuation and B) tip tilt actuation. Figure 5 2 1 Transient response s of TTP mirror arrays. A) Device Type A and B) Type B B A B A Piston resonance Tip tilt resonance

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148 5.4.3 Phase Control Experiment To demonstrate the phase control capability by the mirrors piston motion, a Fraunhofer diffraction experiment is performed using two adjacent sub-mirrors on the mirror array (Type B) Fig ure 5 2 2 illustrates the schematic of the experiment Figure 5 2 2 Schematics of phase control experiments A) Classical Fraun hofer diffraction by two mirror s and B) phase control of one mirror. A He Ne laser light source is first expanded by a 10 objective lens collimated by a collimating lens and then illuminate d on the whole mirror array. i is the incident angle of the light beam and d d enotes the diffraction angle T he classical diffracted amplitude by two rectangular mirrors (in X direction) can be written as : [181] dx ikx y E Ew w i d 2 /2 / 0)) sin (sin exp( ) ( (5 1 ) And can be further derived as: ) cos( 2 ) ( sin )) sin (sin 2 cos( 2 )) sin (sin 2 ( sin0 0 c d E kd ka c d E Ei d i d (5 2 ) where z x x ka kai d i d2 ) ( ) sin (sin 2 z x x kd kdi d i d2 ) ( ) sin (sin 2 X i d ix dx 1x 2 / d 2x 2 / a 2 / w 0 Piston control A B z

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149 In our experiment, as shown in Figure 5 21 B), a small tilt angle between the two mirrors is needed to form the coincidence of the two diffracted beam on the screen, resulting in an additional phase shift t The refore, to recover the classical diffraction pattern, a phase compensation p by the mir rors piston motion is needed. In this case, Equation 5 1 is rewritten as following, with p t : dx z x x kx i E dx z x x kx i E Ea x a x i d a x a x i d 2 / 2 / 0 2 / 2 / 02 2 1 1)] ( exp[ )] ( exp[ dx z x x kx i E z x x kx i z xx kx ia a i d i d i d 2 / 2 / 0 2 1)] ( exp[)]} ( exp[ )] ( {exp[ ) ( sin )]} sin( ) [sin( )] cos( ) {[cos(0 c i d E (5 3 ) So that the intensity: ) ( sin } )] sin( ) [sin( )] cos( ) {[cos(22 2 c I (5 4 ) Figure 5 2 3 Intensity profile s of the diffraction pattern s by phase control of the TTP mirror array. A) C alculated and B) measured results with 2 N C) calculated and D) measured results with ) 1 2 ( N and E) a comparison of bot h experimental and theoretical results with 2 N A C B D E

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150 The classical diffraction pattern can be recovered only at 2 N in which N is an integer Fig ure 5 2 3 shows the calculated results (in two dimensi ons) of both 2 N and ) 1 2 ( N with the corresponding experimental results obtained by controlling the p hase of one mirror. 5. 5 Large -Aperture MEMS Mirror for Free -Space Optical Communications Free -space optics (FSO) has been extensively investigated for applications in space (inter satellite) and terrestrial (inter buildings) optical communications. Compared to optical fibers, FSO is much more versatile, and it is portable and can be easily installed in cities where insta llation of optical fibers is almost impossible by cost and logistics [182, 183] Compared to RF commun ications, FSO is more secure and has much higher bandwidth and no spectrum license issues. There are a few challenges to FSO laser communications. One of the challenges is the beam divergence in free space. In order to maintain high fidelity point to poin t communication, the laser spot size needs to be small even after the laser beam propagates a long distance. -mrad ranges are desired for space and terrestrial applications, respectively. Since the beam divergence angle is inversely proportional to the optical aperture size, large transmitter apertures, typically in at least centimeter ranges, are required FSO systems using electrostatically actuated MEMS mirrors and corner -cube retroreflectors have been reported [184186] However, their aperture sizes are limited to the sub -millimeter range and require complicated fabrication and delicate assembling. Electrothermal bimorph actuation, on the other hand, has the capability of achieving large

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151 aperture and large scan range simultaneously. In this section electothermally actuated MEMS mirror design with large aperture size up to 10 mm is presented 5.5.1 Large -Aperture Mirror Design and Fabrication The MEMS mirrors are based on a lateral -shift -free (LSF) large -vertical displacement (LVD) electrothermal actuator design previously presented. As shown in Fig ure 5 2 4 the mirror plate is symmetrically supported by a series of LSF LVD actuators at two opposite edges. The initial elevation of t he mirror plate is over 0.5 mm, which provides sufficient room for large tiptilt scan motions even for large mirror plates. The initial elevation is due to the residual stress of the bimorphs after device release. Figure 5 2 4 SEMs of fabricated large aperture micromirrors. A -B) LSFLVD actuators C) a 6 mm MEMS mirror, and D ) a 10 -mm MEMS mirror 5 mm D 1 mm A B Al/Pt/ SiO 2 Frames 5 mm C

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152 Fig ure 5 2 4 A ) and B) show the details of the LSF LVD actuator structure which is composed of five folded segments including three Al/SiO2 bimorph beams with embedded Pt heaters and two Si -supported rigid frames. The lengths of the five segments are properly designed to ensure zero lateral shift and tilting for the mirror plate during vertical actuation. A vertical motion and a tip ti lt motion can be obtained by simultaneously and alternatively driving the opposite actuators, respectively. The devices are fabricated by the same process in section 5.4.1 coated) and the rigid frames. Two types of devices have been fabricated with mirror sizes of 6 mm and 10 mm, supported by 56 and 96 actuators respectively. The quality of the mirror surface has been measured by a Wyko NT100 white light profilometer; the radius curvature is about 0.69 m and the average roughness is about 40 nm. 5.5.2 Devic e Characterization The devices are characterized in terms of DC and frequency resp onse. Figure 5 2 5 shows the measured DC responses of both 6 -mm and 10-mm mirrors. Over half millimeter vertical scan can be obtained by simultaneously driving the opposite actuators at 7 V as shown in Figure 5 25 A) Figure 5 25-B) shows the 1 D optical s can angles above 15 and 10 are achieved by the tip tilt mode for the 6 -mm and 10 -mm mirrors, at power of 879 mW and 1.44 W respectively Figure 5 25 C) and D) are the microscopic pictures of a 10-mm mirror being actuated at different tip -tilt position. The frequency responses of the tiptilt mode are measured by a position-sensitive detector with swept frequencies at applied voltage of 3Vdc 0.5Vac As shown in Figure 5 26, two resonance peaks at 178 Hz and 306 Hz for a 6 -mm mirror, and 148 Hz and 234 Hz for a 10 -mm mirror are observed. Both devices have a transient response of about 30 ms (90% rising time)

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153 Figure 5 2 5 Measured DC response s of large aperture micromirrors. A ) V ertical displacement versus applied voltage simu ltaneously on all actuators, B) tip tilt angle versus applied volta ge on actuators at each side, C) and D ): microscopic pictures of a 10 mm mirror at different tip tilt a ngles Figure 5 2 6 Measured frequency responses of both 6 -mm and 10 -mm large -aperture mirrors B C D A

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154 An optical experiment has been performed to characterize the beam propagation profile. A He Ne laser, expanded using a 10 objective lens, is incident on the MEMS mirrors with a beam size of about 5 mm. A collimating lens between the MEMS mirror and the objective lens is used to compensate the curvature of the MEMS mirror and collimate the outgoing beam. Figure 5 27 compares the results of a 1 -mm and 6 -mm MEMS mirror both of which have the same radius curvature. As show n in Figure 5 27, the output beam from the 1 mm mirror diverges much faster than the 6 mm one. The measured divergence angles are respectively about 2.1 mrad and 0.53 mrad for the 1 -mm and 6 -mm mirrors. The 10 -mm mirror has a divergence similar to t hat of the 6 -mm mirror due to the 5 -mm incident beam size Figure 5 2 7 Pictures and intensity profiles of the reflected b eams (incident beam size ~5 mm) by micromirrors with different aperture sizes. A ) A 1 -m m micro mirror and B) a 6 -mm micro mirror at distances of 2 m ( i ), 3.5 m ( ii ), and 7 m ( iii), respectively ( i ) ( ii ) ( iii ) B ( i ) ( ii ) ( iii ) A

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155 5. 6 MEMS -based Miniature Fourier Transform Spectrometer 5. 6 .1 Fourier Transform Spectrometer (FT S ) System Infrared spectr oscopy has been a mature technique for material analysis over 70 years of development [187] It is well known that each material has its unique IR absorption peaks that correspond to the vibration frequencies between bonds of atoms making up the material T herefore the spectrum of a broadband i nfrared light passing through a sample represents the fingerprints of every different kind of material in the sample. In addition, the absorption peak values directly indicate the amount of materials present. T his makes IR spectroscopy an excellent tool fo r quantitative analysis. The early infrared spectroscopy was mainly based on dispersive or filter methods which use a monochromator to produce an infrared spectrum one pixel element at a time [187] Based on a Michelson interferometer, Fourier transform infrared spectroscopy (FTIR) was first int roduced by Fellgett [188] in 1949, but it wa s not until the late 1960s when microcomputers were cap able of performing fast Fourier transform that FTIR s pectroscopy became popular in research labs [189] Figure 5 28. A simplified sch ematic of an FTIR spectrometer ( s ample chamber doesn t show) Figure 5 28 is a simplified schematic that explains the FTIR working principle. It s basically a Michelson interferometer with a movable mirror in one optical path Light fr om the M ova ble mirror Fixed mirror Light source Beamsplitter Detector

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156 source is split into two beams by the beam splitter: one reflected to a fixed mirror and one to the movable mirror which moves around the zero path difference (ZPD) point and generates the optical path difference The beams interfere when they meet back at the beam splitter. A s the moving mirror scans back and forth, the detector observes the cosine interferogram with bright and dark bands corresponding to partial and total constructive and destructive interference in the detected output beam. The in frared light source is p olychromatic with a broad range of wavelengths, so the result ed interferogram represents the sum of all cosine waves generated by each individual infrared wavelength. T he frequency and intensity of each cosine wave in the interferog ram can be resolved by performing t he F ourier transform which converts the measured intensity -versus -mirror -displacement signal i,e., an interferogra m, into a plot of intensity versus wavelength, i.e., a spectrum. FTIR spectro scopy is much more effectiv e and preferred over the dispersive or filter methods for infrared spectral analysis. A major advantage of FTIR is the multiplex advantage (Fellgett's a dvantage ), that is, every data point in the interferogram containing the information of each frequency i n the entire band so it is much more efficient than separat ing and measuring individual frequency as in the dispersive method Other advantages include the through put advantage (Jacquinot s advantage) due to the enhanced light gathering capability by the l arger aperture size instrument, f requency p recision (Conne's a dvantage), and speed advantage etc. [187] High resolution and high sensitivity as well as much faster measur ing speed can be easily achieved. In addition, the setup is relatively simple since there is only one moving part. FTIR spec trometer s ha ve been widely used as a highly effective tool of spectral and material analysis in areas of chemical analysis environmental safety, biohazard detection, planetary geology and space exploration [190193]

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157 5. 6 .2 MEMS -based M iniatu re FT S The conventional FT S systems are bulky and expensive due to the macro-sized instruments used for the tra n slatory scanning a nd precise position monitoring and tilting control of the movable mirror. For instance, the worlds smallest commercial FTIR spectrometer claimed by the Bruker Optics [194] still measures about 22 cm by 30 cm in space (about 8 by 11) and weighs about 7 kg (about 13 lbs). This limits the current FTIR analysis mostly on a bench top e xperimental set up. There are compelling needs for low cost, compact and portable FTIR system, for example, the realization of small portable sensor solutions for industry applications such as color measurement and in -situ process control. Also for space a pplications minimizing system weight and size is very crucial. The key component of an FTIR system that limits its miniaturization is the macro -sized scanning mechanism used for the movable mirror. MEMS technology has provided great potential for device an d system miniaturization in a variety of areas. Researches investigating the FTIR system miniaturization by using MEMS technology have also begun recently [195198] For instan ce, Manzarda et al. demonstrated a translatory moving micromirror driven by push -pull electrostatic actuators that can generate an 80 displacement with a spectral resolution of 5.2 nm at 633 nm achieved [195] Solf et al. reported an electromagnetic driven mirror with a scanning range of about 54 and obtained a 24.5 nm resolution at 1544 nm [196] A silicon optical bench FTIR system recently reported by Yu et al. [197] used an electrostatic comb drive actuator similar to [195] achieved a maximum mirror scanning range of about 50 with a 50 nm resolution for 1580 nm input wavelength. Ho wever, there is a common resolution limitation involved in the above designs. The minimum spectral resolution of a FTIR spectrometer is approximately given by L /2 where is the input wavelength and L is the maximu m optical path difference (OPD) generated by the movable

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158 mirror [187] The resolution can also be equivalently defined by wavenumber to make it independent of input wavelength; in this case, the minimum resolution in wavenumber is approximately equal to 1/L As can be seen in the above mentioned work, the spectral resolution is strongly limited by the small achievable OPD, i.e., by the small scan range of the actuators used for driving the MEMS mirrors. Larger scan range of by electrostatic actuator with about 25 cm1 resolution has been reported more recently in [198] but the device needs to be operated in vacuum which sacrifices the cost -efficiency and compactness of the system. 5. 6 .3 LSF -LVD MEMS mirror for Miniature FT S As demonstrated in Chapter 3, a MEMS mirror with large vertical displacements of over half a millimeter has been realized by a novel LSF LVD electrothermal actuator design. This MEMS mirror has great potential to be used as the translatory scanning mirror for a miniature FT S system A new LSF LVD actuator design which is the same as for the glass lens scanner in s ection 5.3 is used for the large translatory scan MEMS mirror. As shown in Figure 5 29, the LSF LVD actuator consists of two rigid frame s and three Al/SiO2 bimorph beams connected in series for tilting and lateral shift compensation The mirror plate (Al coating on a Si layer) is supported by two LVD actuators at the opposite edges Each actuator consists of two pairs of LSF LVD actuator s that connect with each other by a rigid frame and shar es the third bimorph beams The third bimorph beams connect all a long the mirror edges to minimize the tilting during the vertical actuation The bimorph beams curl upward after release resulting in an initially -elevated mirror rest position With electrothermal actuation on the embedded Pt heater the bimorphs bend down ward due to different thermal expansion of Al and SiO2 and a net vertical displacement of the mirror plate can be obtained The MEMS mirror is fabricated using the same process as for

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159 the lens scanner Figure 5 29 shows an SEM of a fabricated device with a mirror size of 1.8 mm1.6 m m Figure 5 29. SEM pictures of LVD micromirror. A ) LVD micromirror B) the LSF -LVD actuator and C) a packaged device with a US dime 5. 6 4 Device Characterization and FTS Experiment The MEMS mirro r demonstrated a maximum vertical displacement about 1mm at 4 V DC voltage applied on both actuators as in Figure 5 30 A ). A maximum tilting angle (MTA) of about 2.5 has been observed The tilting is mainly due to the fabrication mismatch o f the heater el ectrical resistances and the two actuator structures To minimize the tilting, the two actuators need to be driven separately and a n optimum voltage ratio between the two driving voltages needs to be experimentally characterized A PSD is employed to meas ure the tilting angle. During the mirror actuation, the PSD detects the shift of the reflected light beam that is normally incident on the MEMS mirror (after being reflected by the beam splitter) and the tilting angle can be calculated accordingly. Fig ure 5 31 shows an MTA 1 mm A C B

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160 of 0.55 with a ramp driving voltage of 0.45 1.32 V on both actuators With one actuation voltage being adjusted to 0.45 1.25 V the MTA is minimized to 0.06 Figure 5 30. DC response of LVD micromirror. A ) V ertical displacement and tilting angle versus actuation voltage B) m icroscopic pictures of a packaged device actuated at 0 V ( i ), 2V ( ii ) and 3.5V ( iii). Figure 5 31. Tilting angles measured by a PSD with same voltage on both actuators and minimized tilting angle at optimum voltage ratio. A B ( ii ) ( iii ) ( i ) X Y

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161 Figure 5 3 2 MEMS -based FTS set up A ) Simplified block diagram of a FTS system, B) e xperimental set up of the MEMS mirror based FTS. LS: Light source. BS: Beam splitter. PD: Photo detector. FM: Fixed mirror. MM: MEMS mirror. Figure 5 33. Interferogram of a He Ne laser by actuation voltages of 0.451.25 V and 0.451.32 V on the two actu ators respectively. Insets showing the mirror s deceleration and acceleration range and a uniform mirror velocity range. The experimental set up of the MEMS mirror based FTS is shown in Figure 5 32. A He Ne laser is used for the spectroscopy measurement experiment Figure5 33 shows the interferograms obtained by ramp waveforms of 0.45 1.25 V and 0.45 1.32 V at 0.5 Hz on two actuators respectively The obtained interferogram gives the interference signal in time domain However, direct correlation of th e interferogram in time domain to the optical path difference (OPD) is not applicable due to the non uniformity of the mirror velocity for the entire actuation range LS PD FM MM BS B G(k) k Display PD z FM MM BS LS Sample I (z) z FFT A

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162 Therefore the data has to be re -sampled to account for the scanning non uniformity Phase correction, adaptive digital filtering and polynomial interpolation methods have been used previously to resample the data linearly with OPD [199, 200] A slightly different approach is used for the non uniform mirror velocity calibration The raw interferogram data is first over sampled using interpolation The fringe maxima and minima are then found, and equally spaced sa mple points are selected between each fringe maximum and minimum Since the spacing of the consecutive points between each maximum and minimum is equal, the non-uniformity of the mirror velocity is corrected by this method. These sample points can be used to calibrate the spectrometer, by using a light source with known wavelength, e.g., a He Ne laser The calibration results can be applied to an interferogram for unknown wavelength by re -sampling it with the calibration sample points As shown in Figure 5 34, spectral resolutions of 71.4 cm1 and 19.2 cm1 were obtained by MEMS mirr respectively. Table 5 3 lists the mirror scan ranges and the resulting spectral resolutions by different actuation voltages. Figure 5 34. S pectr a of He Ne las er obtained by MEMS -based FTS with different spectral resolutions A ) 71.4 cm1 and B) 19.2 cm1. (Note: voltages listed in Table 5 -3) A B

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163 Table 5 3 MEMS mirror scan ranges, tilting angles and spectral resolutions at different driving voltages Actuat ion voltage (V) Maximum tilting angle (degree) Mirror scan ra nge Spectral resolution (cm1) 0.45 1.32 0.55 0.45 1.25/1.32 0.06 70 71.4 0.45 2.25 1.41 0.45 2.04/2.25 0.22 261 19.2 5. 6 5 Mirror-Tilt -Insensitive FTS System The conventional FTS system is very sensitive to the mirror tilting during its tra nslatory scan. For MEMS mirrors, tilting during large range scanning is almost inevitable As presented in the previous section, although the MEMS mirror has large translatory scan range close to 1 mm, the useful range with good interference signal s is onl y limited to about 2One possible solution is to modify the FTS setup and make it less sensitive to the tilting of MEMS mirror. A novel mirror tilt insensitive (MTI) FTS using the dual reflective LVD MEMS mirror and a corner -cube retroreflector is proposed to compensate the small tilting of the MEMS mirror. 5.6.5.1 MTI -FTS setup Fig ure 5 35 shows the MTI FTS concept The core of the system is a Michelson interferometer in which both light beams from the beam splitter are directed to the dual reflective MEMS mirror There are one corner -cube retroreflector and one fixed mirror in the two arms of the interferometer, respectively At the rest position, the MEMS mirror is placed at the zero OPD point With the actuation of the MEMS mirror the OPD of the two sub arms can be generated and interferograms can be detected by a photodetector The tilting of the scanning mirror is compensated by taking the advantages of the dual reflections of the MEMS mirror and the retroreflector whose reflected light beam is always parallel to the incident beam When there is a tilting of the MEMS mirror in any direction, both light beams will bounce off the MEMS mirror with the equal but opposite tilting Due to the right angle arrangement of the corner -cube, the

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164 returning light beam will be shift ed to the opposite direction. Therefore the two returning light beams will have the same amount of shift at the same direction when propagating back to the beam splitter Note that the coincident beams do move laterally on the photodetector. So the active area of the photodetector should be sufficient ly large. A simulation for a 1 tilt is shown in B), which verifies this tilt compensation concept T herefore the interferogram detected by the photodetector will not be affected by the mirror tilting Figure 5 35. MTI -FTS system A ) S chematic of the MTI -FTS set up, B) s imulation showing the tilting compensation (by FRED) C) a photo of the MTI -FTS demonstration setup. CCR: Corner -cube retroreflector. PD: Photodetector. BS: Beam spl itter. MM: MEMS micromirror. FM: Fixed mirror. LS: Light source. MM (1 tilt) LS B CCR BS FM PD CCR FM MM (zero tilt) BS LS PD A CCR BS MM FM LS P D C

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165 Another advantage of this MTI FTS is that the equivalent OPD is four times o f the physical scan range of the MEMS mirror, which is doubled compared to a conventional FTS so that the spectral resolutio n can be further enhanced Figure 5 35 C) shows a picture of the experimental setup that measures about 1255 cm3. S maller dimensions can be obtained by further reducing the sizes of the components such as the beam splitter and retroreflector. 5. 6.5.2 Experimental results T he same fabrication process presented in section 4.2 .2 is used to achieve the dual reflective surface s for the MEMS mirror. The MEMS mirror has been characterized in terms of vertical displacements and tilting angle s at both DC and AC actuation, and similar results as the original device in subsection 5.6.4 are obtained. As shown in Fig ure 5 36 A ), a maximum vertical displacement over 1 mm has been obtained at small driving voltages of less than 5 V with a good linear actuation r ange of about 0.6 mm (from 1.2 V to 3 V) Figure 5 36. Dual reflective LVD MEMS mirror characterization A ) DC response of vertical displacement and tiltin g angle versus applied voltage and B ) AC response of tilting angles in two directions A maximum tilting angle of 1.7 has been observed in X direction (as labeled in Figure 5 30) The tilting angle is mainly due to the fabrication mismatch on the two actuators such as the A Driving voltage Tilting in X Tilting in Y B

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166 thickness of bimorph beams, nonuniformity of Si undercut of the frames and the mirror edges The tilting angle during AC actuation (as in Figure 5 36 B)) is also characterized by using the MTI -FTS setup with the photodetector replaced by a position-sensitive detector (PSD) The PSD detects the lateral shif t of the light beams reflected by the MEMS mirror and the tilting angle can be calculated accordingly. FTS experiments were performed using a He Ne laser on a conventional Michelson interferometer based FTS and the MTI FTS by the same MEMS mirror The MEM S mirror was scanning vertically by 0.4 mm at 0.1 Hz Figure 5 3 7 A ) shows the interferogram obtained from a conventional FTS, which was greatly impaired by the tilting of the MEMS mirror The interference signal is detected only at small driving voltages of each driving cycle, where the tilting angle is very small. As a comparison, the MTI -FTS is much less sensitive to the mirror tilting and gives good interference signals in the entire cycle by the same mirror actuation as in Figure 5 3 7 B). Figure 5 37. Interferograms of the He -Ne laser A) by a c onventional FTS setup, B) by a MTI FTS setup with micromirror actuation C) at a ramp voltage from 1.2V to 2. 5 V at 0.1Hz. C B A

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167 Figure 5 3 8 Spectr a of He Ne l aser obtained by the MTI FTS system with MEMS mirror being actuated at different ramp waveform s A ) at 1.2V to 2V and B) at 1.2V to 2.5V respectively Inset compares the resolutions of 19.2 cm1 and 8.1 cm1. Figure 5 38 shows the spectra of a He Ne laser obtained at different mirror scan ranges after the calibration of nonuniform mirror scan Spectral resolutions of 19.2 cm1 and 8. 1 cm1 were obtained by the mirror actuation ranges of about 131 08 OPD of 522 3 mm, respectively. 5 7 Summary This chapter present s the applications of the developed MEMS optical scanners. 2 D scanning endoscope designs based on the new LSF -LVD micromirror for nonlinear opti cal and OCT imaging are developed and in vivo 3 D images are successfully obtained The potentials in other applications have also been investigated. Glass lens scanners by a modified LSF LVD actuator design demonstrate large vertical scan range of about 0.9 mm with small tilting angle less than 0.4. TTP mirror arrays with about 200 motion and 36 tiptilt scan with a fill -factor of 65% are presented for optical phased array applications. An LSF LVD based MEMS mirror with an aperture size as large as 10 mm demonstrates the great potential for free space optical communication s Finally, prototype s of novel miniature FT S systems utilizing the A B

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168 LSF -LVD based MEMS mirror are demonstrated A novel mirror tilt insensitive FTS system is combined with a dual reflective MEMS mirror and high spectrum resolution up to 8.1 cm1 is achieve d.

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169 CHAPTER 6 CONCLUSION AND FUTURE WORK As presented in Chapter 1, this research effort has been following the main objective of th e research project -the development of new MEMS optical scanners and the MEMS -based endoscopes for biomedical imaging applications. First, a novel LSF LVD elect rothermal actuator design is presented. The new design successfully solves the problems of large lateral shift, large initial tilting angle, and low fill factor of previous hinged -bimorph designs Versatile optica l scanners for in vivo biomedical imaging based on the novel actuator design have been developed, including a high-fill -factor TTP MEMS mirror for 2 D lateral scan and large vertical lens scanners for axial scan. I maging endoscopes employing the 2 D MEMS m irrors have been developed and 3 -D in vivo imaging results are successfully obtained for both nonlinear optical and OCT imaging. Second, a new full -circumferential scanning ( FCS ) imaging endoscope design is proposed and developed for internal organ imaging applications. The FCS endoscope uses dual reflective 1 D mirror that can be fabricated by a novel s elf aligned deep trench process. A dual -foldedbimorph ( DFB ) actuator is designed and the fabricated device successfully solves a series of problems of the hinged-bimorph design. I t has achieved fast thermal response that enables f ull circumferential scan range at real time imaging speed ; the mirror plate has zero i nitial tilting and zero initial elevation as well as a stationary rotation axis along the mirro r center. A similar self aligned deep trench process is also developed for light -weight large aperture MEMS mirrors. Last, the versatile MEMS optical scanners based on the new actuator designs open up great opportunities in many other applications, which are also explored in this research effort. Such investigations include the TTP mirror array for optical phased array applications, large aperture

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170 size MEMS mirror for free -space optical communications, and the large vertical scan MEMS mirror for miniature FTS systems. 6. 1 Research Accomplishments The following tasks have been accomplished in this research effort : 1 A novel LSF -LVD actuator has been designed and it solves the large lateral shift and large initial tilting angle problem in the previous LVD a nd translatory optical scanners. 2 V ersatile optical scanners have been fabricated based on this new LSF LVD actuator design. TTP tilt scan above 60 at small driving voltages less than 5 V. O ptimized thermal design achieves fast thermal response with cut -off frequency up to 78 Hz. 3 LSF -LVD based lens scanners with integrated polymer and glass lens es are demonstrated The glass lens scanner achieves about 0.8 mm vertical scan of the focal plane at the driving voltage smaller than 5 V, with small tilting angle less than 0.4 4 A 5 mm diameter imaging endoscope for nonlinear optical imaging is developed utilizing the LSFLVD based 2 D MEMS mirror with a high fill -factor of 25% 3 -D n onlinear optic al imaging with is achieved. 5 A 5 .8 mm diameter imaging endoscope using the same 2 D MEMS mirror for a time domain OCT system is developed. In vivo imaging with 3 D reconstruction of a 2.3 2.3 0 .6 mm3 vol ume at the axial and lateral resolutions about 10 and 6 are demonstrated. 6 A novel self aligned deep trench process is developed and dual -reflective MEMS mirrors are fabrica ted for FCS imaging. A similar process is also used to fabricate light -weight and largeaperture MEMS mirrors. Device with a mass reduction up to 82.3% is fabricated and significant improvement on the devices mechanical stiffness has been achieved. 7 A novel DFB actuator is designed for FCS mirrors. The fabricated device demonstrates over 90 mechanical rotation angle at real time imaging speed above 60 Hz. The new design solves the problems of large initial mirror tilting and elevation, and achieves a stationary rotation axis along the mirror center. 8 4 TTP mirror arrays have been tip tilt scan and fill -factor of 65% with large subaperture size of 0.9 mm are obtained, the potential for optical phased array applications are also demonstrated. 9 LSF -LVD based MEMS mirrors with aperture size as la rge as 10 mm for free -space optical communications are presented Piston scan above 0.5 mm and 1-D scan above

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171 10 optical angles with resonance in the order of 100 Hz are achieved. Due to the large aperture size, divergence angles as small as 0.5 mrad have been obtained 10. P rototype s of novel miniature FT S systems utilizing the LSF LVD based MEMS mirror are demonstrated A novel mirror tilt insensitive FTS system is combined with a dual reflective MEMS mirror and high spectrum resolution up to 8.1 cm1 is a chieved. 6.2 Future Work The versatile MEMS optical scanners and imaging endoscopes developed in this research effort have realized the intended in vivo biomedical endoscopoic imaging applications and demonstrated great potentials in other applications a s well The following research efforts are suggested for future work. 1 For reliable clinical usage of the MEMS -based imaging endoscope s the device reliability such as long term reliability and repeatability, and sustainability to external shock needs to be studied Experiments and optimization on actuator structure s and material and process variations are suggested for the investigation 2 The LVD lens scanners and FCS MEMS mirrors are ready for the integration into imaging endoscopes, which need to be devel oped for confocal microscopy, OCM and FCS imaging applications. 3 The fabricated l ight -weight MEMS mirrors have poor surface quality due to the thermal stress of the single layer Al thin film. Multi layer thin film structure with st r ess compensation need s to be investigated to improve the mirror surface quality. 4 The fill -factor of the TTP mirror arrays is still limited. A n ew process needs to be developed to hide the actuators therefore to achieve higher fill factor. 5 Large aperture MEMS mirrors need f urther investigation s for larger range and more versatile scan capability Experiments are needed to verify their application for free space optical communications. 6 To further verify the miniature FTS applications, experiments with broadband light sources need to be investigated and more efficient data processing methods need to be developed. Compact and portable packages are also needed.

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172 APPENDIX PUBLICATIONS AND PROVISONAL PATENTS GENERATED BY THIS RE SEARCH EFFORT The following is a list of archival journal (J), conference ( C) publications and provisional patents (P) generated from the accomplished research effort. Full -Circumferential Scanning (FCS) MEMS mirror s : J 1 : L. Wu and H. Xie, A 124 d egre e r otation a ngle e lectrothermal m icromirror with i ntegrated p latinum h eater ," IEEE Journal of Selected Topics in Quantum Electronics, v ol 13 2, pp 316 321, Apr. 2007. J2: L. Wu and H. Xie, An electrothermal micromirror with dual reflective surfaces f or circumferential scanning endoscopic imaging, Journal of Micro/nanolithography, MEMS and MOEMS, Vol. 8, No.1, 013030, 2009. C1: L. Wu and H. Xie, "A large rotation angle electrothermal micromirror with integrated p latinum h eater," presented at the 2006 I EEE/LEOS International Conference on Optical MEMS, Big Sky, MT, Aug. 2006. C 2 : K.S. Lee, L. Wu, H. Xie, O. Ilegbusi, M. Costa and J.P. Rolland, "A 5 -mm catheter for constant resolution probing in Fourier domain optical coherence endoscopy," presented at th e Photonics West 2007, San Jose, Mar. 2007. C 3 : L. Wu and H. Xie, A dual reflective e lectrothermal MEMS m icromirror for f ull circumferential s canning e ndoscopic i maging, presented at the Photonics West 2008, San Jose, Jan. 2008. C 4 : L. Wu and H. Xie, "A scanning micromirror with stationary rotation axis and dual reflective surfaces for 360 forward-view endoscopic imaging," presented at the 1 5 th IEEE International Conference on Solid -State Sensors, Actuators and Microsystems (Transducers '0 9 ), Denver, CO, Ju n 2009. Lateral -Shift Free (LSF) Large -Vertical -Displacement (LVD) actuator: J 3 : L. Wu and H. Xie, "A large vertical displacement electrothermal bimorph microactuator with very small lateral shift," Sensors and Actuators A, vol 145 146, pp: 371379, Jul -Aug. 2008. C 5 :L. Wu and H. Xie, A large vertical displacement electrothermal bimorph micromirror/lens," presented at the 14th IEEE International Conference on Solid State Sensors, Actuators and Microsystems (Transducers '07) Lyon, France, Jun. 2007. P1 : L. Wu and H. Xie, Electrothermal microactuator for large vertical displacement without tilt or lateral shift (provisional, UF 612P)

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173 LSF -LVD lens scanner s : C 6 : L. Wu and H. Xie, "A l ateral -shift -free LVD m icrolens s canner for c onfocal m icroscopy," presen ted at the 2007 IEEE/LEOS International Conference on Optical MEMS, Hualien, Taiwan, Aug. 2007. C7 : L. Wu and H. Xie, A tunable microlens with 0.9 mm scan range and small lateral shift, presented at the 2009 IEEE/LEOS International Conference on Optical MEMS & Nanophotonics Clearwater Beach FL, Aug. 2009. Light w eight MEMS mirror : P2: L. Wu and H. Xie, A self aligned deep trench process for light -weight MEMS micromirror ( provisional, UF # 12565) Nonlinear optical imaging endoscope: C 8 : L. Wu, L. Fu, A. Ja in, T. Nishida, M. Gu, and H. Xie, "An e ndoscopic n onlinear o ptical i maging p robe b ased on 2 -D micr omirror," presented at the 2007 IEEE/LEOS Annual Meeting, Lake Buena Vista, FL, Oct. 2007. C 9 : L. Fu, L. Wu, S. Russell, H. Xie, M. Gu, 3 D t issue i maging w ith a n onlinear o ptical e ndoscope p robe, presented at Focus on Microscopy 2007, Valencia, Spain, Apr. 2007. C 10: D. Morrish, L. Wu, H. Xie, J. Reynolds, A. Boussioutas and M. Gu, Nonlinear imaging by an endoscope probe incorporating a TTP microelectrome chanical system mirror. presented at OSA Spring Optics & Photonics Congress, Vancouver, BC, Canada, Apr. 2009. OCT imaging endoscope: C 1 1 : S. Guo, L. Wu, J. Sun, L. Liu and H. Xie, "Three dimensional optical coherence tomography based on a high -fill -facto r microelectromechanical mirror presented at OSA Spring Optics & Photonics Congress, Vancouver, BC, Canada, Apr. 2009. C1 2 : S.Guo, A.Pozzi, H.Ling, J. Sun, L. Wu, L. Liu and H. Xie, 3 D polarization sensitive optical coherence tomography of canine menis cus based on a 2 D h igh -fill f actor m icroelectromechanical mirror, presented at the 31st Annual International Conference of the IEEE Engineering in Medicine and Biology Society (EMBC'09), Minneapolis, Minnesota, Sep. 2009. C1 3 : K. Jia, S. Pal, L. Wu D. H amilton and H. Xie, Dental optical coherence tomography employing miniaturized MEMS -based imaging probe presented at the 2009 IEEE/LEOS International Conference on Optical MEMS & Nanophotonics Clearwater Beach FL, Aug. 2009. Tip -tilt -piston mirror arrays :

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174 C 14: L. Wu, S.B. Maley, S.R. Dooley, T.R. Nelson, P.F. MacManamon and H. Xie, "A large aperture, piston-tip tilt micromirror for optical phase array applications," presented at the 21 th IEEE International Conference on Micro Electro Mechanical Systems (MEMS '0 8 ), Tucson, AZ, Jan. 2008. P3: H. Xie, S.B. Maley, P.F. Macmanamon, T.R. Nelson, L. Wu, K. Jia and A. Pais, Desi gn and fabrication method for high -fill -factor micromirrors/micromirror arrays (provisional, UF 710P) Large -aperture MEMS mirror for free -space optical communications : C1 5 : L. Wu and H. Xie, Large aperture, rapid scanning MEMS micromirrors for free space optical communications presented at the 2009 IEEE/LEOS International Conference on Optical MEMS & Nanophotonics Clearwater Beach F L, Aug. 2009. MEMS -based miniature FTS systems : C1 6 : L. Wu, A. Pais, S.R. Samuelson, S. Guo and H. Xie, "A miniature Fourier transform spectrometer by a large -vertical -displacement microelectromechanical mirror ," presented at OSA Spring Optics & Photonics Congress, Vancouver, BC, Canada, Apr. 2009. C1 7 : L. Wu, A. Pais, S.R. Samuelson, S. Guo and H. Xie, "A mirror -tilt -insensitive Fourier transform spectrometer based on a large vertical displacement micromirror with dual reflective surface," presented at the 1 5 th IEEE International Conference on Solid -State Sensors, Actuators and Microsystems (Transducers '0 9 ), Denver, CO, Ju n 2009. P4: H. Xie, L. Wu, A. Pais, S.R. Samuelson and S. Guo, MEMS -b ased FTIR s pectrometer (provisional, UF 738P) P5: L. Wu, S. Guo, A. Pais, S.R. Samuelson and H. Xie, Mirror -t ilt i nsensitive Fourier t ransform s pectrometer (provisional, UF 777P)

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191 BIOGRAPHICAL SKETCH Lei Wu w as born in Dangtu, Anhui, China in December of 1980. He enrolled in University of Science and Technology of China, Hefei, Anhui, China in the fall of 1998. Based on his academic excellence, he was awarded with early graduation and received his B.S. deg ree in Physics in June 2002, and with exemption of examinations he enroll ed in the graduate program in Shanghai Institute of Microsystems and Information Technology (SIMIT), Chinese Academy of Sciences, Shanghai, China. During the three years of M.S. study at SIMIT, Lei was introduced to the MEMS area, and focused on developing micro-hotplate for MEMS gas sensors and bolometers. He also worked on MEMS device testing and developed a real time readout interface for MEMS infrared sensor arrays. Lei received his M .S. degree in Microelectronics and Solid State Electronics in June of 2005. Lei joined Biophotonics and Microsystems Lab, Interdisciplinary Micrsosystems Group at University of Florida in August of 2005 to pursue his Ph.D. degree, and started his research developing MEMS optical scanners for endoscopic biomedical imaging applications. His research interests include MEMS actuators, optical scanners, endoscopoic biomedical imaging, photonics, optical phased array, free -space optical communications and MEMS ba sed miniature Fourier transform spectroscopy. Upon the completion of his dissertation, he has contributed 20 research publications and 5 provisional patents from his doctoral research. Lei is a member of the Institute of Electrical and Electronics Enginee rs and the Optical Society of America, and is also a member of the Eta Kappa Nu honor societies. He pursue d a career in the area of optical MEMS after receiving his Ph.D. degree in 2009.